DESIGN AND ANALYSIS OF A MEMS VIBRATION SENSOR FOR AUTOMOTIVE MECHANICAL

SYSTEMS

BY

Joel Rebello

A thesis submitted in conformity with the requirements

for the degree of Master of Applied Science

Graduate Department of Mechanical and Industrial Engineering

University of Toronto

copy Copyright by Joel Rebello (2009)

ABSTRACT

Design and Analysis of a MEMS Vibration Sensor for Automotive Mechanical Systems

Joel Rebello

Master of Applied Science

Graduate Department of Mechanical and Industrial Engineering

University of Toronto

2009

This thesis presents the theoretical analysis and experiment results of MEMS sensors

designed for the application of low frequency vibration sensing Each sensor consists of a

proof mass connected to a folded beam micro-flexure with an attached capacitive comb

drive for displacement sensing Three comb drive arrangements are evaluated the

transverse lateral and tri-plate differential The sensors are fabricated using the well

developed foundry processes of PolyMUMPS and SoiMUMPS In addition a capacitance

to voltage readout circuit is fabricated using discrete components Static tests evaluating

the capacitance to displacement relation are conducted on a six degree of freedom

robotic manipulator and dynamic tests evaluating the sensor response to sinusoidal

excitations are conducted on a vibrating beam The end use of the sensor involves real-

time vibration monitoring of automobile mechanical systems such as power seats

windshield wipers mirrors trunks and windows allowing for early detection of

mechanical faults before catastrophic failure

ii

ACKNOWLEDGMENT

I would first like to thank God for all his blessings and express my gratitude to my

parents and brother for their endless support love and encouragement

I would like to thank my supervisors Dr William Cleghorn and Dr James Mills for

their guidance I would also like to thank my friends and colleagues Henry Chu Dr

Pezhman Hassanpour Shael Markin Dr Lidai Wang and Dr Xuping Zhang at the

Laboratory for Nonlinear Systems Control for great conversation guidance and

friendship Thank you to the members of my Examination Committee Dr Foued Ben

Amara and Dr Jean Zu for their time and feedback

Finally I would like to extend my appreciation to CMC Microsystems for the MEMS

device fabrication and design support and to Autorsquo21 a Network of Federal Centers of

Excellence for project funding

iii

TABLE OF CONTENTS 1 INTRODUCTION 1

11 RESEARCH PURPOSE 1 12 MICROELECTROMECHANICAL SENSORS 2 13 AUTOMOTIVE SENSORS 3 14 OPERATING ENVIRONMENT 4 15 OBJECTIVES AND CONTRIBUTION 5 16 RESEARCH OUTLINE 5

2 LITERATURE REVIEW 7

21 EXISTING METHODS FOR FAULT DETECTION 7 211 Motor Current Monitoring 7 212 Temperature Sensors 8 213 Acoustic Emission Sensors 9

22 MEMS VIBRATION SENSORS 10 221 Resonant Sensors 10 222 Piezoelectric Sensors 11 223 Displacement Variation Sensors 12

3 SENSOR STRUCTURE 15

31 SENSING 15 32 MECHANICAL ANALYSIS 21

321 Micro-Flexure Selection 21 322 Equations of Motion 26 323 Quality Factor 30

33 DRIVE STABILITY 32 34 ELECTROSTATIC ACTUATION 34 35 SECTION SUMMARY 35

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT 36

5 NOISE ANALYSIS 40

51 MECHANICAL-THERMAL NOISE 40 52 ELECTRICAL NOISE 41

6 FABRICATION OVERVIEW 43

61 POLYMUMPS 43 62 SOIMUMPS 44

7 EXPERIMENTAL RESULTS 46

71 MEMS SENSOR ndash T1 46 72 READOUT CIRCUITRY 51 73 MEMS SENSORndash L1 52 74 SECTION SUMMARY 59

8 CONCLUSION AND FUTURE WORK 61

81 CONCLUSION 61 82 FUTURE WORK 62

iv

LIST OF TABLES TABLE 3-1 SPECIFICATIONS FOR EACH CAPACITIVE COMB DRIVE IMPLEMENTED IN THIS WORK T1 L1 AND

T2 21 TABLE 3-2 DIMENSIONS AND MECHANICAL PROPERTIES FOR THE THREE MEMS VIBRATION SENSORS

OUTLINED IN THIS WORK 30 TABLE 3-3 SUMMARY OF THE THEORETICAL COEFFICIENTS OF DAMPING AND QUALITY FACTORS FOR THE

THREE SENSORS PRESENTED IN THIS WORK 32 TABLE 3-4 SUMMARY OF THE CHARACTERISTICS OF THE SENSORS OUTLINED IN THIS WORK 35 TABLE 7-1 SUMMARY OF EXPERIMENTAL RESULTS 60

v

LIST OF FIGURES

FIGURE 3-1 AN INTER-DIGITATED COMB DRIVE WHERE CHANGES IN CAPACITANCE ARE GENERATED BY

EITHER CHANGES IN GAP DISTANCE X0 OR IN THE OVERLAP AREA Y0timesT 16 FIGURE 3-2 AN INTER-DIGITATED COMB DRIVE IN A DIFFERENTIAL TRI-PLATE ARRANGEMENT WHERE X01

X02 AND X03 ARE THE GAP DISTANCES AND X01= X02ltlt X03 TO IMPROVE SENSITIVITY 18 FIGURE 3-3 COMPARISON OF THE EFFECT OF GAP DISTANCES ON LINEARITY FOR TWO CASE lsquorsquo REPRESENTS

CASE 1 WHERE THE FINGERS IN ARE IMPLEMENTED IN A DIFFERENTIAL TRI-PLATE ARRANGEMENT AND X01= X02= 2 ΜM AND X03= 15 ΜM lsquorsquo REPRESENTS THE CASE 2 WHERE THE FINGERS ARE IMPLEMENTED IN A SINGLE ENDED ARRANGEMENT AND X01= X02=X03=2 ΜM 19

FIGURE 3-4 lsquorsquo REPRESENTS THE OUTPUT OF THE DIFFERENTIAL TRI-PLATE DRIVE WHILE lsquorsquo REPRESENTS THE OUTPUT OF THE TRANSVERSE DRIVE THE DASHED LINE REPRESENTS THE LINEAR APPROXIMATION FOR DISPLACEMENTS LESS THAT 1 ΜM FOR THE PURPOSE OF COMPARISON THE OUTPUT OF THE TRANSVERSE DRIVE WAS EQUALIZED TO ZERO AT ZERO DISPLACEMENT 20

FIGURE 3-5 SENSOR STRUCTURE 22 FIGURE 3-6 CLAMPED-CLAMPED FLEXURE 24 FIGURE 3-7 CRAB LEG FLEXURE 25 FIGURE 3-8 FOLDED BEAM FLEXURE 25 FIGURE 3-9 COMPARISON OF THE NON-LINEAR DEFLECTION AMONG THE THREE MICRO-FLEXURES

CONSIDERED CLAMPED-CLAMPED CRAB-LEG AND FOLDED BEAM ALL FLEXURES WERE DESIGNED TO HAVE THE SAME STIFFNESS OF 035 N M-1 26

FIGURE 3-10 LUMPED MASS APPROXIMATION WITH BASE EXCITATION YB(T) 27 FIGURE 3-11 LUMPED MASS APPROXIMATION WITH DIRECT MASS EXCITATION Z(T) 28 FIGURE 3-12 GRAPH REPRESENTING STABILITY REGIONS FOR GAP CLOSING CAPACITIVE DRIVES lsquorsquo

REPRESENTS THE RHS OF EQ (32) FOR T2 WHILE lsquorsquo REPRESENTS THE RHS OF T1 34 FIGURE 4-1 IDEAL CHARGE AMPLIFIER 36 FIGURE 4-2 CAPACITANCE TO VOLTAGE READOUT CIRCUIT WHICH IS MADE UP OF FOUR ELEMENTS THE

CHARGE AMPLIFIER VOLTAGE AMPLIFIER DEMODULATOR AND LOW PASS FILTER 37 FIGURE 4-3 ORCAD SIMULATION RESULTS OF THE CAPACITANCE TO VOLTAGE READOUT CIRCUIT THE

EXCITATION VOLTAGE VE IS 75 VPK AND THE OUTPUT OF THE VOLTAGE AMPLIFIER VVA IS APPROXIMATELY 13125 VPK 38

FIGURE 4-4 ORCAD SIMULATION RESULTS SHOWING THE SINUSOIDAL OUTPUT OF THE DEMODULATOR VDM WITH A 984 VPK-PK AMPLITUDE AND THE DC OUTPUT OF THE LOW PASS FILTER VLP AT APPROXIMATELY 5 V 39

FIGURE 5-1 ORCAD SIMULATION OF THE NOISE VOLTAGE SPECTRAL DENSITY ENO AT THE OUTPUT OF THE VOLTAGE AMPLIFIER 42

FIGURE 5-2 TOTAL OUTPUT NOISE ENO AT THE VOLTAGE AMPLIFIER WHICH IS EXPRESSED AS ENO=( int ENO2)12

42 FIGURE 6-1 POLYMUMPS FABRICATION PROCESS FOR A SIMPLE CANTILEVER BEAM 44 FIGURE 6-2 SOIMUMPS FABRICATION PROCESS FOR A SIMPLE CANTILEVER BEAM 45 FIGURE 7-1 THE SOLID LINE REPRESENTS THE THEORETICAL CAPACITANCE VERSUS DISPLACEMENT

RELATION WHILE THE SOLID POINTS REPRESENT THE EXPERIMENTAL VALUES THE DASHED LINE REPRESENTS THE ASSUMED LINEAR RELATION FOR DISPLACEMENTS LESS THAN 1ΜM THE EXPERIMENTAL AVERAGE DRIVE SENSITIVITY IS APPROXIMATELY 0024 PF ΜM-1 FOR DISPLACEMENTS GREATER THAN 12 ΜM DRIVE INSTABILITY IS OBSERVED AND IS ATTRIBUTED TO THE PULL-IN EFFECT 47

FIGURE 7-2 lsquotimesrsquo REPRESENTS THE AMPLITUDE AND SOLID LINE REPRESENTS THE PHASE RESPONSE THE RESONANCE PEAK IS LOCATED AT 3600 HZ AND ACCOMPANIED BY A PHASE SHIFT OF 168deg 48

FIGURE 7-3 A SEM MICRO-GRAPH SHOWING THE SENSOR TI FABRICATED USING POLYMUMPS THE FOUR FOLDED BEAM FLEXURES HAD A LENGTH OF 650 ΜM AND TOTAL STIFFNESS OF 14 NM THE EFFECTIVE PROOF MASS WAS 29E-9 KG AND THE SYSTEM HAD A NATURAL FREQUENCY OF 3500 HZ 50

FIGURE 7-4 A SEM MICRO-GRAPH OF THE TRANSVERSE COMB DRIVE IMPLEMENTED IN T1 WITH 200 COMB FINGERS A GAP DISTANCE OF 2 ΜM OVERLAP LENGTH OF 45 ΜM AND THICKNESS OF 2 ΜM THE

vi

AVERAGE EXPERIMENTAL SENSITIVITY WAS FOUND TO BE 0024 PF ΜM-1 FOR DISPLACEMENTS OF LESS THAN 1 ΜM 50

FIGURE 7-5 THE SOLID LINE REPRESENTS THE THEORETICAL CAPACITANCE TO VOLTAGE RELATION EXPRESSED BY EQ (37) WITH A SENSITIVITY OF 141 V PF-1 THE SOLID POINTS REPRESENT THE EXPERIMENTAL DATA POINTS WHEN DISCRETE CAPACITORS 1-4 PF ARE USED IN THE READOUT CIRCUIT THE HORIZONTAL DASHED LINE REPRESENTS THE OUTPUT VOLTAGE OF THE MEMS SENSOR AT REST AND CORRESPONDS TO A CAPACITANCE OF 35 PF 52

FIGURE 7-6 OSCILLOSCOPE TRACE OF THE INPUT EXCITATION VOLTAGE VE asymp 75 VPK AND OUTPUT OF THE VOLTAGE AMPLIFIER VVA asymp 103 VPK READINGS WERE TAKEN WITH THE MEMS SENSOR AT REST 53

FIGURE 7-7 OSCILLOSCOPE TRACE OF THE DEMODULATOR OUTPUT VDM asymp 76 VPK-PK AND OUTPUT OF THE LOW PASS FILTER VLP asymp 36 V READINGS WERE TAKEN WITH THE MEMS SENSOR AT REST 53

FIGURE 7-8 lsquorsquo REPRESENTS THE EXPERIMENTAL DATA POINTS WITH THE DASHED LINE REPRESENTING THE FITTED CURVE SOLID LINE REPRESENTS THE THEORETICAL DATA POINTS ACCORDING TO EQ (10) 54

FIGURE 7-9 EXPERIMENT SETUP WHICH INCLUDES THE MEMS SENOR AND READOUT BOARD MOUNTED ON THE CLAMPED-FREE BEAM AND TEST EQUIPMENT 55

FIGURE 7-10 CLOSE-UP OF THE MEMS SENSOR AND READOUT BOARD MOUNTED ON THE TEST BEAM ALONG WITH THE COMMERCIAL PIEZOELECTRIC ACCELEROMETER ONLY THE CHARGE AMPLIFIER (CA) AND THE VOLTAGE AMPLIFIER (VA) ARE MOUNTED ON THE SAME PROTOTYPE BOARD AS THE MEMS SENSOR DUE TO SPACE CONSTRAINTS 55

FIGURE 7-11 FIGURE COMPARING THE RESPONSE OF A COMMERCIAL PIEZOELECTRIC ACCELEROMETER AND MEMS SENSOR (L1) TO A BEAM VIBRATION A) DYNAMIC RESPONSE OF THE PIEZOELECTRIC ACCELEROMETER REPRESENTING THE TRUE BEAM VIBRATION WITH A PEAK VOLTAGE 114 V CORRESPONDING TO AN APPLIED ACCELERATION OF 114 G B) DYNAMIC RESPONSE OF THE MEMS SENSOR WITH A PEAK VOLTAGE OF 006 V 56

FIGURE 7-12 SENSOR DYNAMIC RESPONSE TO THE VIBRATION OF THE CLAMPED-FREE BEAM AND REPRESENTS A DIRECT EXPORT FROM THE DYNAMIC SIGNAL ANALYZER (AGILENT 35670) TOP TRACE IS THE TIME DOMAIN SIGNAL OF THE LOW PASS FILTER OUTPUT WHERE VLPMAXasymp73MV WHICH REPRESENTS A MAXIMUM ACCELERATION OF APPROXIMATELY 15G THE BOTTOM TRACE REPRESENTS THE FREQUENCY SPECTRUM WITH A PEAK AT 23HZ WHICH CORRESPONDS TO THE BEAM NATURAL FREQUENCY 57

FIGURE 7-13 EXPERIMENTAL RESULTS SHOWING THE RESPONSE OF THE SENSOR TO VARIOUS ACCELERATIONS IN THE RANGE OF 0-20 G WHERE lsquorsquo REPRESENTS THE EXPERIMENTAL DATA POINTS AND THE SOLID LINE REPRESENTS THE FITTED CURVE THE EXPERIMENTAL SENSITIVITY WAS FOUND TO BE 4258 MV G-1 58

vii

NOMENCLATURE a applied acceleration A applied acceleration amplitude App parallel plate overlapping area Asldie-film effective area for slide-film damping c coefficient of damping Cf feedback capacitance Cpp capacitance of the parallel plate capacitor cslide-film coefficient of damping as a result of slide-film damping

csqueeze-film coefficient of damping as a result of squeeze-film damping CTotal total capacitance of a comb drive CTotalBottom Total capacitance of the bottom drive in a tri-plate comb drive CTotalL total capacitance of a lateral drive CTotalT total capacitance of a transverse drive CTotalTop Total capacitance of the top drive in a tri-plate comb drive d distance between the parallel plates E Youngrsquos Modulus of sensor material eno noise voltage spectral density Eno total output noise voltage EnoAMP total output amplifier noise voltage EnoR total output resistor noise voltage

f excitation frequency in Hz fn system natural frequency in Hz

Fact electrostatic actuation force

FN noise force Fx-Elect total electrostatic force for pull-in instability Itruss second moment of inertia of truss section on the folded beam flexure Ibeam second moment of inertia of beam section on the folded beam

flexure kair dielectric constant of air at atmospheric pressure kB Boltzmannrsquos constant kT total stiffness of parallel folded beam flexure kx stiffness of various flexures in x-direction ky stiffness of various flexures in y-direction L flexure beam length Lplate effective plate length

viii

Ls shin beam length Lt thigh beam length m effective proof mass N number of comb fingers on the movable comb

NB number of balls

Qtotal total quality factor

Qf energy lost to surrounding fluid

Qs energy lost through supports

Qm energy lost through sensor material S theoretical sensitivity of sensor SL sensitivity of a lateral comb drive SM theoretical mechanical sensitivity of the proof massflexure system SRO sensitivity of readout circuit ST sensitivity of a transverse comb drive STAVG average sensitivity of a transverse comb drive t time T absolute temperature tb flexure beam thickness tc out of plane comb finger thickness

Vact electrostatic actuation voltage VCA output amplitude of charge amplifier VDM output amplitide of demodulator VE input amplitude of sinusoidal excitation voltage VLP output of low pass filter VLPMAX maximum output voltage at low pass filter VVA output amplitude of voltage amplifier vy relative motion of proof mass with respect to base Vy relative motion amplitude w flexure beam width Wplate effective plate width ws shin beam width wt thigh beam width x0 initial comb finger gap

x01 small gap distance on top side of the tri-plate comb drive

x02 small gap distance on bottom side of the tri-plate comb drive

ix

x03 larger gap distance on the tri-plate comb drive

XN noise displacement y0 initial comb finger overlap length yB base displacement YB base displacement amplitude ym absolute proof mass displacement α parameter dependant on the WplateLplate ratio ∆C Change in capacitance from a tri-plate drive in a differential

arrangement ε permittivity of free space κ second moment of inertia ratio (shinthigh) ζ damping ratio microair absolute viscosity of air ω base excitation frequency in rads ωn sensor natural frequency in rads ωv electrostatic actuation voltage frequency in rads L1 Sensor fabricated using SoiMUMPS with a lateral comb drive T1 Sensor fabricated using PolyMUMPS with a transverse comb drive T2 Sensor fabricated using SoiMUMPS with a tri-plate comb drive

x

1 INTRODUCTION

11 RESEARCH PURPOSE

The modern automobile consists of many mechanical systems such as power seats

windshield wipers mirrors trunks and windows which are all susceptible to breakdown

Without any condition monitoring system the breakdown is usually catastrophic and

requires an expensive part replacement Real-time condition monitoring allows for early

detection of faults which could require a simple solution such as the application of a

lubricant to fix This prolongs the useful life of the component and prevents sudden and

unexpected failure Real-time condition monitoring can be accomplished by examining the

vibration signature of a mechanical system For example an automobile power window

consists of a DC motor and its associated bearings and couplings a gear reduction system

consisting of worm and spur gears and kinematic links Faults resulting in excessive

vibrations may be caused by coupling misalignment bearing failure or gear train failure

Coupling misalignments occur at the connection between the drive shaft and the driven

shaft and are typically due to imperfect manufacturing Bearing failure is caused by lack

of lubrication or moisture contamination causing rusting while gear train failure is caused

by misaligned broken cracked or chipped gear teeth [1] Each fault occurs at a

characteristic frequency and so the state of the mechanical system can be determined by

monitoring the amplitudes of the relevant frequencies Vibrations due to coupling

misalignments occur at harmonics of the shaft rotational speed Gear vibrations occur at

the gear turn speed or at sidebands of the gear mesh frequency [1] Ball bearing vibrations

1

2

may be caused by outer bearing race defects fOB inner bearing race defects fIB ball

defects fB and train defects fT which all occur at specific frequencies [12]

cos( )12

dBOB r

p

bNf fd

β = minus

cos( )12

dBIB r

p

bNf fd

β = +

2cos( )1

2p r d

Bd p

d f bfb d

β = minus

cos( )12

drT

p

bffd

β = minus

(1)

where NB is the number of balls fr is the rotation frequency bd is the ball diameter β is the

contact angle between the ball and races and dp is the ball pitch diameter By monitoring

the real time vibration signature of the mechanical system anomalies can be quickly

identified and fixed

12 MICROELECTROMECHANICAL SENSORS

Microelectromechanical systems (MEMS) inertial sensors provide a small footprint

with sensitivities that are either comparable or exceed any macro sensor along with the

capability of mass production and low unit cost These sensors utilize compliant micro-

flexures attached to a proof mass that displaces in response to an environmental

acceleration Many transduction mechanisms have been developed that convert the

displacement into a measurable electric signal and include thermal piezoresistive

piezoelectric optical and capacitive methods Most MEMS sensors are silicon based and

are fabricated using surface or bulk micromachining Surface micromachining creates free

standing movable structures on top of a substrate using a combination of sacrificial layers

3

and structural layers which are commonly polysilicon [3] In bulk micromachining the

mechanical structures are defined using a removal process where bulk material typically

silicon is etched away MEMS sensors have been used for vibration and shock monitoring

on industrial systems and robotics guidance and navigation in global positioning systems

(GPS) seismometry in earthquake prediction image stabilization in digital cameras and

automobile safety and stability [4]

13 AUTOMOTIVE SENSORS

Sensors cover every major aspect of a modern automobile power-train sensors

monitor fuel combustion and emissions chassis sensors monitor road traction and tire

condition and body sensors facilitate air-bag deployment and vehicle proximity for radar

guided cruise control [5] Pressure sensors which typically consist of a piezoresistive

strain sensing element attached to a silicon diaphragm that deflects when exposed to an

applied pressure are one of the first micro-machined sensors used in an automobile

Implemented as manifold absolute pressure (MAP) sensors they allow precise control of

the air fuel ratio which allows the catalytic converter to efficiently reduce tailpipe

emissions [6] Variable reluctance sensors based on electromagnetics are used for

automobile traction control and produce a voltage output that is dependant on the

magnetic flux variations between a rotating component and the sensor bias magnet [57]

MEMS based linear accelerometers are utilized for airbag deployment upon impact and

provide a lower cost and smaller package over the older ball-in-tube accelerometers [6]

4

14 OPERATING ENVIRONMENT

Through proper packaging MEMS sensors can withstand the harsh automotive

environment where they face cyclic temperatures (-40degC to 150degC) mechanical shock or

vibration (10g-500g) and exposure to various fluids (oil brake fluid or engine coolant)

and airborne particles [8] Temperature changes induce thermal stresses causing

undesirable bending of the silicon microstructures that changes their dynamic response

Mechanical shock which could occur during a high impact collision or vibration from

uneven road conditions cause the delicate microstructures to fracture at points of stress

concentration These effects could be reduced by efficient design Use of a folded beam

micro-flexure that allows expansion and contraction or use of fillets at corners which

reduce the chance of fracture Bonding pads wire bonding and onboard circuit

components can be protected from air borne particles which could cause shorting using

parylene or gels Alternatively the entire package can be sealed hermetically which also

protects the MEMS device [8] In some cases such as fluid pressure sensors intimate

contact between the MEMS device and potentially corrosive fluid is necessary A recent

approach used a nanometer thin coating of silicon carbide (SiC) over a foundry fabricated

MEMS sensor using low temperature low pressure chemical vapor deposition (LPCVD)

[9] The SiC known for its chemical inertness did not compromise the mechanical

characteristics of the original design Therefore MEMS provides a strong platform for an

automotive sensor

5

15 OBJECTIVES AND CONTRIBUTION

The objective of this work is the development of a low frequency MEMS vibration

sensor The sensor must be low cost and mass producible which can be achieved using

standard foundry fabrication processes such as the Polysilicon Multi-User MEMS

Processes (PolyMUMPS) or the Silicon-on-Insulator Multi-User MEMS Processes

(SoiMUMPS) In addition the number of additional processing steps which are required

if integrated CMOS circuitry is used or if vacuum packaging is required must be

minimized This can be achieved using an off-chip circuit made with widely available

discrete components and sensor operation in air

In addition this work presents a MEMS sensor that is fabricated using standard

foundry processes such as PolyMUMPS and SoiMUMPS compared to many MEMS

sensors in literature that are fabricated using customized processes

16 RESEARCH OUTLINE

This work demonstrates three iterations of a low frequency capacitive vibration

sensor fabricated using the standard foundry processes Each sensor consists of a

suspended proof mass that displaces in response to an external vibration The proof mass

is attached to a folded beam micro-flexure compliant in one direction and its lateral

movement is sensed using a capacitive comb drive The first sensor T1 is fabricated in

PolyMUMPS and uses a transverse capacitive comb drive The second L1 fabricated in

SoiMUMPS uses a fully linear lateral capacitive drive and achieves an increase in

sensitivity from T1 Finally the third sensor T2 is fabricated in SoiMUMPS and uses a

differential tri-plate capacitive drive This sensor takes advantage of the increase in

6

sensitivity gained from a transverse drive but rejects off-axis displacement errors and

common mode using a differential tri-plate arrangement In addition an off-chip

capacitance to voltage readout circuit is fabricated and tested on sensor L1

Chapter 2 begins with an overview on the various methods of fault detection

followed by a review of MEMS vibration sensors each with a different transduction

mechanism Chapter 3 outlines the sensor structure focusing on the mechanical design

and sensing principal Chapter 4 focuses on the readout circuitry layout and simulation

results Chapter 5 outlines the various source of noise including mechanical noise arising

from the sensor and electric noise from the readout circuit Chapter 6 overviews the

foundry fabrication processes used including PolyMUMPs and SoiMUMPS The

experimental results of two sensors T1 and L1 which were fabricated and tested within

the thesis time frame are presented in Chapter 7 Chapter 8 concludes the thesis with a

brief summary and outline for future work

2 LITERATURE REVIEW

21 EXISTING METHODS FOR FAULT DETECTION

As this thesis develops a sensor for fault detection using vibration monitoring this

section overviews three existing methods for fault detection in mechanical systems First

motor current monitoring is discussed as it has the ability to measure significant bearing

faults in motor systems Next temperature sensors specifically thin film thermocouples

(TFTC) and MEMS temperature sensors are discussed These sensors are mounted in the

chamber of a lubricating fluid or in close proximity to a moving component and monitor

temperature fluctuations caused by increased wear Finally acoustic emission (AE)

sensors are discussed which have the ability to detect ultrasonic emissions generated by

plastic deformation in the crystalline lattice of a mechanical component

211 MOTOR CURRENT MONITORING

Monitoring of stator current is used as a method of determining bearing related faults

that produce radial motion between rotor and stator This motion varies the air gap flux

density producing stator currents at predictable frequencies fbng that represent the

bearing faults [10]

bng e vf f m f= plusmn sdot (2)

where fe is the electrical supply frequency m=123hellip and fv is one of the characteristics

frequencies listed in the equations of (1)The bearing damage must be significant to

produce noticeable current fluctuations and it may be difficult to distinguish from current

fluctuations due to voltage imbalances machine saturation or statorarmature faults [2]

7

8

212 TEMPERATURE SENSORS

Temperature sensors are largely used in high temperature high load and high speed

environments typically turbine engines and industrial tools Thermocouple sensors

consist of two dissimilar metals joined together at two junctions and when exposed to a

temperature difference generate a thermoelectric voltage dependant only on the material

properties and junction temperature difference [11] In turbine engines these sensors can

be placed on bearing casings or in the lubricant to monitor temperature fluctuations due

to increased friction The thermocouples are fabricated using bulk materials or micro-

machined thin films (TFTC) with the latter providing a faster response small package

and substantially lower production costs [12] The TFTC can be embedded for in situ

monitoring of processes in hostile environments One study successfully embedded

TFTC into electroplated nickel and attained device characteristics on par with surface

mounted devices [13] More recently MEMS temperature sensors have been developed

that consist of micro-machined semiconductor material that undergoes structural

deformation with temperature changes A common implementation is a bent beam

structure whose temperature induced deflection δ is expressed as [14]

( ) 3

2 2

sin12 cos sinz

EA T LEI AEL

α βδ 2β β

∆=

+

(3)

Temperature sensors are well suited for localized fault monitoring and detection but can

only characterize a mechanical system consisting of multiple components through the use

of multiple sensors

9

213 ACOUSTIC EMISSION SENSORS

Acoustic emission (AE) sensors have been used to characterize wear in machine

tools and monitor bearing and gear problems in centrifugal pumps [1516] First

developed as a Non-Destructive Testing (NDT) technique to detect cracks in civil

structures these sensors detect acoustic emissions generated by the release of vibration

waves in a crystalline lattice due to plastic deformation or crack growth [16]

Measurements are made using piezoelectric transducers with high natural frequencies

100 kHz to 1 MHz to capture the ultrasonic AE emissions An AE sensor is useful as it

has the ability to detect subsurface cracks in gear teeth or bearings before appearing on

the surface causing further damage [17] More recently MEMS acoustic sensors have

been developed and one design includes multiple transducers on a single substrate

which each detect acoustic emission energy at different frequencies [18] This helps

distinguish spurious acoustic emissions arising from impact and friction from those

arising from plastic deformation When compared to typical piezoelectric sensors the

MEMS devices have lower sensitivities and fail to detect some acoustic emissions

[1819] In addition the acoustic emission signal suffers severe attenuation as is crosses

various interfaces such as a gearbox or bearing casings In one experiment consisting of

a pinion gear and an associated bearing a 44dB attenuation was seen between an AE

sensor placed directly on the pinion to one placed on the bearing casing and in some

cases intermediate loss of the signal was observed [20]

10

22 MEMS VIBRATION SENSORS

This thesis focuses on the development of a MEMS vibration sensor for fault

detection as the vibration signature of a mechanical system has the potential to give an

overview of the entire system Many types of MEMS vibration sensors exist with

different operating principals based on resonance the piezoelectric effect and

displacement variation

221 RESONANT SENSORS

Resonant sensors are based on the resonant frequency shift of a fixed-fixed micro-

beam that is excited into resonance using electrostatic thermal or piezoelectric methods

Under an applied axial load there is a shift in the resonant frequency which is expressed

as [21]

2 2

2

1 12 1

nn n

EI Fl2

fl A Eα γ

π ρ= +

I

(4)

where l is the beam length E is Youngrsquos Modulus I is the moment of inertia ρ is the

density A is the cross section area F is the applied axial force αn and γn are mode

dependant coefficients and are 4730 and 0295 for mode 1 in the case discussed The

axial force F is generated by a proof mass attached to one beam end which moves in

response to an external acceleration The performance of resonant sensors are dependant

on the quality factor Q which is the ratio of total energy stored in the system to the

energy lost per cycle and determines the sharpness and amplitude of the resonance peak

A high Q resonator develops a peak that is easily distinguishable in a phase locked loop

control circuit and allows for improved performance and resolution [3] A resonator Q

value is lowered by energy losses to the surrounding fluid attached supports and

11

material Losses to the fluid surrounding the MEMS structure is the dominant loss

mechanism for MEMS sensors operated in air and can be greatly reduced by vacuum

packaging which can be accomplished using glass-silicon anodic bonding [22] In

addition to increasing fabrication cost the long term stability of the vacuum degrades

with ultimate device failure stemming from leakage through micro cracks and defects

and out-gassing [22] Support losses arise from the restoring forces they generate and

can be reduce by balanced designs such as a double ended tuning fork (DETF) or

operation in higher modes [3] In addition the sensitivity of a resonant sensor can be

increased by the addition of micro-levers to increase the axial force In one such design a

two stage micro-lever increased the sensitivity by an order of magnitude over a more

conventional single lever arm [2324]

222 PIEZOELECTRIC SENSORS

Piezoelectric materials produce a charge in response to an applied force This is an

intrinsic effect in materials such as Zinc Oxide (ZnO) whereas materials such as Barium

Titanate or Lead Zirconate Titanate (PZT) need to be poled by placing the material in a

strong electric field at an elevated temperature [3] The piezoelectric effect is expressed

mathematical by [25]

i ij j ik kD d T Eε= + (5)

where D represents the electrical displacement (Cm2) d is the charge coefficient (CN)

T is the applied stress matrix (Nm2) ε is the permittivity matrix (Fm) and E is the

electric field matrix (Vm) The subscripts i=123 j=123456 and k=123 represent

the direction to which the physical property is related For micro-sensors poled in the out

of plane direction only the D3 term is relevant and if fabricated using ZnO or PZT only

12

the charge coefficients d31 d32 and d33 are relevant [25] The piezoelectric material is

commonly deposited as a thin film on the surface of compliant structures in bulk micro-

machined sensors or incorporated as a layer in surface micro-machining on a similar

structure [26] However the inclusion of multiple materials introduces large temperature

variations causing out of plane bending especially in surface micro-machined devices

[2627] Poled piezoelectric materials could be depolarized which serves to reduce the

piezoelectric affect in the material This can occur if the sensor is exposed to a strong

electric field of the opposite polarity high temperatures in excess of the Curie point or

high mechanical stress Also the piezoelectric effect reduces over time an effect that can

be reduced with the addition of composite elements [28] Foundry processes such as

PolyMUMPS and SoiMUMPS do not incorporate piezoelectric layers thus sensors using

this method of transduction must use a customized process The piezoelectric sensor

outlined in [26] uses polysilicon surface micro-machining similar to PolyMUMPS

however it incorporates a ZnO layer that is not available in the foundry process

223 DISPLACEMENT VARIATION SENSORS

Displacement variation sensors consist of a proof mass connected to a compliant

micro-flexure The movement of the proof mass is sensed using optical electron

tunneling or more commonly capacitive methods A recent optical accelerometer

utilized a novel multilayer nano-grating for in-plane displacement sensing of the proof

mass [29] For small space variations between the nano-gratings a large change in the

optical reflection amplitude of the grating was observed [30] Even though optical

sensors offer resolutions approaching the Brownian noise limit they are not available in

small packages that are easily placed in space constricted areas for a low cost [29]

13

The electron tunneling effect is observed when the proximity between two electrodes

one flexible and the other fixed is on the order of 10Ǻ [31] If the two electrodes are

biased with voltage VB a tunneling current It flows between the gap [31]

exp( )t B I tgI V xαinfin sdot minus Φ (6)

where α1 is a constant and equal to 1025 Ǻ-1 eV-frac12 Ф is the effective height of the

tunneling barrier and xtg is the tunneling gap When the flexible electrode moves in

response to acceleration a feedback controller maintains the original distance while

determining the applied acceleration Electron tunneling sensors are able to sense

accelerations in the microg-ng range due to the exponential relation between tunneling current

and gap However these highly sensitive devices are susceptible to thermo-mechanical

noise which could be reduced by operation at low pressure [32] It is also important for

the two electrodes to be coated with metal which is difficult to achieve with foundry

fabrication

Capacitive sensors are based on the parallel plate capacitor and are implemented

using inter-digitated comb fingers in an in-plane lateral or transverse drive or in an out of

plane parallel plate drive [33] These sensors have the advantage of high sensitivity low

power consumption small package low temperature dependence and are easily

integrated with existing foundry fabrication processes [34] The lateral drive offers a fully

linear response but low sensitivity whereas the transverse and parallel plate drives

achieve high sensitivity if gap distances are of a few microns and large capacitance areas

are utilized However both drives have a non-linear response and are prone to pull-in

Also if these sensors are fabricated using a surface micromachining process such as

PolyMUMPS where the ground plane is a few microns above the structural layer they

14

are prone to electrical failure due to shorting This was observed in an experiment where

similar sensors were placed in a vibrating environment with a peak of 120g and shorting

between the structural layer and ground plane was found to be the predominate mode of

failure [35] However with certain fabrication processes such as SoiMUMPS this issue

can be neglected as the substrate material below any moving component is removed

Most capacitance sensors are fabricated using an integrated CMOS (Complementary

Metal Oxide Semiconductor)-MEMS technique where the MEMS sensor is fabricated

before during or after the CMOS circuit fabrication [36] Although these sensors offer

high-sensitivity and small parasitic capacitance they suffer from in-plane and out of

plane curling of the beam which reduces the capacitance between adjacent comb fingers

[37] In addition multiple processing steps are required for device fabrication

More than the other transduction methods MEMS capacitive displacement sensors are

attractive for practical implementation in an automobile and therefore is the selected

method for the vibration sensor presented in this work Their low temperature dependence

is an ideal characteristic for an automobile environment that faces varying temperature If

implemented with off-chip readout circuitry they require a standard foundry fabrication

process which provides a strong platform for mass production and low unit cost

3 SENSOR STRUCTURE

This chapter presents the physical structure of the three MEMS vibration sensors

which includes the displacement sensing capacitive comb drive and the micro-

flexureproof mass system First the capacitive comb drive and the three variations

implemented the transverse drive lateral drive and differential tri-plate drive are

discussed Next the mechanical structure is presented and includes the micro-flexure

selection quality factor and equations of motion Finally instability that places a limit

on the operating range of the sensor as a result of a gap closing capacitive comb drive is

discussed

31 SENSING

Capacitive sensors are based on the parallel plate capacitor and are commonly

implemented using inter-digitated comb fingers A capacitance can be realized if a voltage

is placed between two closely spaced plates If the fringing electric field is neglected the

capacitance of the parallel plate capacitor Cpp (F) is expressed as [28]

air pppp

k AC

dε

= (7)

where kair=1 is the dielectric constant of air at atmospheric pressure ε=88542times10-12 F m-1

is the permittivity of free space App is the overlapping area (m2) and d is the distance

between the two plates (m)

15

16

Figure 3-1 An inter-digitated comb drive where changes in capacitance are generated by either changes in gap distance x0 or in the overlap area y0timest

Figure 3-1

For a set of inter-digitated comb fingers or comb drive shown in the total

capacitance Ctotal is expressed as [28]

0 0

0 0

( ) ( )( ) (Total cy y y yC N t

)x x x xε

+ += + + minus

(8)

where N is the number of comb fingers on the movable comb y0 is the initial comb finger

overlap length (m) tc is the out of plane comb finger thickness (m) and x0 is the inital

finger gap (m) In this work the comb drive is implemented in three arrangements

transverse lateral and differential tri-plate In a transverse comb drive changes in

capacitance are generated by transverse movements along the x-axis resulting in changes

to the gap x0 between the movable and fixed comb fingers The nonlinear relation

between total capacitance CTotalT and displacement is expressed as

17

00 0

1 1( ) (Total TC N ty

)x x x xε

= + minus +

(9)

Although the drive sensitivity part is not constant an average sensitivity can be

assumed for small displacements [38]

Total TC xpart

In a lateral comb drive changes in capacitance are generated by lateral movements

along the y-axis resulting in changes to the overlapping area y0timestc The total capacitance

CTotalL between all comb fingers is expressed as

00

2 ( )Total LN tC yx

yε= +

(10)

The sensitivity of the lateral comb drive SL is expressed as

0

2Total LL

C N tSy x

εpart= =

part

(11)

The lateral comb drive was first introduced by W C Tang T H Nguyen and R T

Howe [33] in an electrostatic resonator and has since been used numerous times in both

drive and sense configurations It has the advantage of a fully linear response and

negligible instability but offers low sensitivity The sensitivity for the lateral drive L1

implemented in this work is 0038 pF microm-1

A differential tri-plate drive shown in Figure 3-2 which is a variation of the

transverse drive is implemented to improve sensitivity and reduce noise [39]

18

Figure 3-2 An inter-digitated comb drive in a differential tri-plate arrangement where x01 x02 and x03 are the gap distances and x01= x02ltlt x03 to improve sensitivity

In the tri-plate drive the at rest gap distances between the two capacitors formed by any

movable finger is not equivalent For a differential arrangement insuring that x01=

x02ltltx03 the linearity for small displacements (less than 1 microm) is enhanced over the case

of a single ended arrangement where x01=x02=x03=x0 which is equivalent to the

transverse drive This is shown graphically in Figure 3-3 where x01= x02=x03=2 microm in

the lower curve and x01= x02= 2 microm and x03= 15 microm in the upper curve A substantial

improvement in linearity is attained with the tri-plate drive

19

0

005

01

015

02

025

03

035

0 02 04 06 08 1

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-3 Comparison of the effect of gap distances on linearity for two case lsquorsquo represents case 1 where the fingers in are implemented in a differential tri-plate arrangement and x01= x02= 2 microm and x03= 15 microm lsquorsquo represents the case 2 where the fingers are implemented in a single ended arrangement and x01= x02=x03=2 microm

The tri-plate drive is also implemented in a differential arrangement where the output

is the change in capacitance rather than the total capacitance resulting from an external

excitation This reduces errors cause by off-axis excitation and common mode noise The

output of the differential tri-plate drive is expressed as [39]

Total Top Total BottomC C C∆ = minus where

003 01

1 1( ) (Total TopC N ty

)x x x xε

= +

minus +

(12)

20

002 03

1 1( ) (Total BottomC N ty

)x x x xε

= +

minus +

A graphic comparison of the capacitancedisplacement relation for sensors T1 and T2 is

shown in Figure 3-4 For these gap closing drives it is assumed that for displacements

less that 1 microm the relation is linear indicated by the dashed line above each curve The

sensitivity within this region is approximately 0046 pF microm-1 and 0253 pF microm-1 for T1

and T2 respectively The order of magnitude improvement is attributed to the differential

tri-plate arrangement and thicker comb finger

0

02

04

06

08

1

12

14

16

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-4 lsquorsquo represents the output of the differential tri-plate drive while lsquorsquo represents the output of the transverse drive The dashed line represents the linear approximation for displacements less that 1 microm For the purpose of comparison the output of the transverse drive was equalized to zero at zero displacement

21

In this work the transverse drive is first implemented in sensor T1 using the

PolyMUMPS fabrication process Then in an effort to linearize the sensor response the

lateral drive is used in L1 using SoiMUMPS The reduction in drive sensitivity due to the

lateral drive is further reduced by increasing the gap distance to 3 microm which is

recommended by the foundry to ensure a successful device These reductions are

balanced by increasing the number of comb fingers N and selecting a substrate thickness

that is 125 times large than T1 The third device T2 is fabricated using SoiMUMPS and

uses the tri-plate differential drive This design capitalizes on the increase in sensitivity of

a gap closing drive and larger substrate thickness attaining an order of magnitude

increase in sensitivity over T1 and L1 The specifications for each drive along with the

successive improvement in drive sensitivity are outlined in Table 3-1

Table 3-1 Specifications for each capacitive comb drive implemented in this work T1 L1 and T2

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Comb Finger Thickness tc (microm) 2 25 25 Gap Distance x0 (microm) 2 3 2 Number of Fingers N 200 256 30 Overlap length y0 (microm) 45 10 70 Drive Sensitivity (pF microm-1) 0046 0038 0273

32 MECHANICAL ANALYSIS

321 MICRO-FLEXURE SELECTION

The sensor structure consists of a proof mass connected to a compliant micro-flexure

that itself is anchored to the sensor substrate The proof mass displaces in response to an

22

applied acceleration and this displacement is sensed using the capacitive comb drive

attached to the proof mass

Proof mass

Folded beam micro-flexure

Anchor

Figure 3-5 Sensor Structure

This section focuses on micro-flexure selection as it should exhibit a large linear

displacement range and have a high lateral stiffness ratio to reduce cross-axis sensitivity

A number of micro-flexures have been used in MEMS research these include the

clamped-clamped crab-leg and folded beam shown in Figure 3-6-through 3-7

The clamped-clamped flexure exhibits a high stiffness ratio however for large lateral

displacements extensional axial forces buildup within the beams resulting in a non-linear

force-displacement relation [40] Within the linear regime the stiffness of the clamped-

clamped flexure is expressed as [40]

3

3b

yEt wk

L=

(13)

23

where E is the Youngrsquos Modulus of the beam material (N m-2) tb is the beam out of

plane thickness (m) w is the beam width (m) and L is the beam length (m) The stiffness

ratio is expressed as [40]

2x

y

k Lk w

=

(14)

If L is two orders of magnitude greater than w this results in a ratio of approximately

10000 The linear displacement range is determined using Finite Element Analysis (FEA)

of a clamped-clamped flexure with L=387 microm w=4 microm and tb=2 microm The result of this

simulation is plotted in and compared to a structure with constant stiffness ky

= 035 N m-1 predicted by Eq (13) The linear relation is valid for small deflections that

are less than 2 microm In addition this micro-flexure is susceptible to buckling as a result of

a compressive residual stress that may arise due to the fabrication process [41]

Figure 3-9

In the crab leg design the thigh section is meant to relieve the extensional axial

stress postponing the occurrence of non-linearity until greater lateral displacement [42]

The stiffness of the crab leg flexure is expressed as [43]

3

44

b s s ty

s s t

Et w L LkL L L

κκ

+= +

(15)

where κ=IsIt=(wswt)3 is a ratio of the second moment of inertia of the shin (s) and thigh

(t) ws and wt represent the width and Ls and Lt represent the length of the shin and thigh

regions respectively The stiffness ratio is expressed as [43]

34

4t s s tx

y s t s

w L L Lkk w L L L

αα t

+= +

(16)

The linear displacement range of a crab leg flexure with Ls=285 microm Lt=30 microm ws=wt=3

microm and tb=2 microm is determined using FEA The results are plotted in Figure 3-9 and

compared to a structure of constant stiffness ky=035 N m-1 predicted by Eq (15) The

24

non-linearity is greatly reduced when compared to the clamp-clamped flexure with Eq

accurately predicting the deflection until 4 microm However the flexure is more

susceptible to off-axis deflection as there is an order of magnitude reduction in the

stiffness ratio as compared to the clamped-clamped flexure

(15)

The folded beam flexure stands out as it is designed to exhibit a large linear

displacement range and to relieve residual stress due to the manufacturing process by

allowing the beams to expand and contract along the axial direction [40] It was first used

in an electrostatic resonator and has subsequently been used for electrostatic actuation in

a magnetic hard drives and micro-positioning [3344] If the truss section is rigid

ItrussgtgtIbeam then the stiffness of the flexure is expressed as [4043]

3

32b

yEt wk

L=

(17)

The stiffness ratio is expressed as [40]

2x

y

k Lk w

=

(18)

Figure 3-6 Clamped-clamped flexure

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

ABSTRACT

Design and Analysis of a MEMS Vibration Sensor for Automotive Mechanical Systems

Joel Rebello

Master of Applied Science

Graduate Department of Mechanical and Industrial Engineering

University of Toronto

2009

This thesis presents the theoretical analysis and experiment results of MEMS sensors

designed for the application of low frequency vibration sensing Each sensor consists of a

proof mass connected to a folded beam micro-flexure with an attached capacitive comb

drive for displacement sensing Three comb drive arrangements are evaluated the

transverse lateral and tri-plate differential The sensors are fabricated using the well

developed foundry processes of PolyMUMPS and SoiMUMPS In addition a capacitance

to voltage readout circuit is fabricated using discrete components Static tests evaluating

the capacitance to displacement relation are conducted on a six degree of freedom

robotic manipulator and dynamic tests evaluating the sensor response to sinusoidal

excitations are conducted on a vibrating beam The end use of the sensor involves real-

time vibration monitoring of automobile mechanical systems such as power seats

windshield wipers mirrors trunks and windows allowing for early detection of

mechanical faults before catastrophic failure

ii

ACKNOWLEDGMENT

I would first like to thank God for all his blessings and express my gratitude to my

parents and brother for their endless support love and encouragement

I would like to thank my supervisors Dr William Cleghorn and Dr James Mills for

their guidance I would also like to thank my friends and colleagues Henry Chu Dr

Pezhman Hassanpour Shael Markin Dr Lidai Wang and Dr Xuping Zhang at the

Laboratory for Nonlinear Systems Control for great conversation guidance and

friendship Thank you to the members of my Examination Committee Dr Foued Ben

Amara and Dr Jean Zu for their time and feedback

Finally I would like to extend my appreciation to CMC Microsystems for the MEMS

device fabrication and design support and to Autorsquo21 a Network of Federal Centers of

Excellence for project funding

iii

TABLE OF CONTENTS 1 INTRODUCTION 1

11 RESEARCH PURPOSE 1 12 MICROELECTROMECHANICAL SENSORS 2 13 AUTOMOTIVE SENSORS 3 14 OPERATING ENVIRONMENT 4 15 OBJECTIVES AND CONTRIBUTION 5 16 RESEARCH OUTLINE 5

2 LITERATURE REVIEW 7

21 EXISTING METHODS FOR FAULT DETECTION 7 211 Motor Current Monitoring 7 212 Temperature Sensors 8 213 Acoustic Emission Sensors 9

22 MEMS VIBRATION SENSORS 10 221 Resonant Sensors 10 222 Piezoelectric Sensors 11 223 Displacement Variation Sensors 12

3 SENSOR STRUCTURE 15

31 SENSING 15 32 MECHANICAL ANALYSIS 21

321 Micro-Flexure Selection 21 322 Equations of Motion 26 323 Quality Factor 30

33 DRIVE STABILITY 32 34 ELECTROSTATIC ACTUATION 34 35 SECTION SUMMARY 35

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT 36

5 NOISE ANALYSIS 40

51 MECHANICAL-THERMAL NOISE 40 52 ELECTRICAL NOISE 41

6 FABRICATION OVERVIEW 43

61 POLYMUMPS 43 62 SOIMUMPS 44

7 EXPERIMENTAL RESULTS 46

71 MEMS SENSOR ndash T1 46 72 READOUT CIRCUITRY 51 73 MEMS SENSORndash L1 52 74 SECTION SUMMARY 59

8 CONCLUSION AND FUTURE WORK 61

81 CONCLUSION 61 82 FUTURE WORK 62

iv

LIST OF TABLES TABLE 3-1 SPECIFICATIONS FOR EACH CAPACITIVE COMB DRIVE IMPLEMENTED IN THIS WORK T1 L1 AND

T2 21 TABLE 3-2 DIMENSIONS AND MECHANICAL PROPERTIES FOR THE THREE MEMS VIBRATION SENSORS

OUTLINED IN THIS WORK 30 TABLE 3-3 SUMMARY OF THE THEORETICAL COEFFICIENTS OF DAMPING AND QUALITY FACTORS FOR THE

THREE SENSORS PRESENTED IN THIS WORK 32 TABLE 3-4 SUMMARY OF THE CHARACTERISTICS OF THE SENSORS OUTLINED IN THIS WORK 35 TABLE 7-1 SUMMARY OF EXPERIMENTAL RESULTS 60

v

LIST OF FIGURES

FIGURE 3-1 AN INTER-DIGITATED COMB DRIVE WHERE CHANGES IN CAPACITANCE ARE GENERATED BY

EITHER CHANGES IN GAP DISTANCE X0 OR IN THE OVERLAP AREA Y0timesT 16 FIGURE 3-2 AN INTER-DIGITATED COMB DRIVE IN A DIFFERENTIAL TRI-PLATE ARRANGEMENT WHERE X01

X02 AND X03 ARE THE GAP DISTANCES AND X01= X02ltlt X03 TO IMPROVE SENSITIVITY 18 FIGURE 3-3 COMPARISON OF THE EFFECT OF GAP DISTANCES ON LINEARITY FOR TWO CASE lsquorsquo REPRESENTS

CASE 1 WHERE THE FINGERS IN ARE IMPLEMENTED IN A DIFFERENTIAL TRI-PLATE ARRANGEMENT AND X01= X02= 2 ΜM AND X03= 15 ΜM lsquorsquo REPRESENTS THE CASE 2 WHERE THE FINGERS ARE IMPLEMENTED IN A SINGLE ENDED ARRANGEMENT AND X01= X02=X03=2 ΜM 19

FIGURE 3-4 lsquorsquo REPRESENTS THE OUTPUT OF THE DIFFERENTIAL TRI-PLATE DRIVE WHILE lsquorsquo REPRESENTS THE OUTPUT OF THE TRANSVERSE DRIVE THE DASHED LINE REPRESENTS THE LINEAR APPROXIMATION FOR DISPLACEMENTS LESS THAT 1 ΜM FOR THE PURPOSE OF COMPARISON THE OUTPUT OF THE TRANSVERSE DRIVE WAS EQUALIZED TO ZERO AT ZERO DISPLACEMENT 20

FIGURE 3-5 SENSOR STRUCTURE 22 FIGURE 3-6 CLAMPED-CLAMPED FLEXURE 24 FIGURE 3-7 CRAB LEG FLEXURE 25 FIGURE 3-8 FOLDED BEAM FLEXURE 25 FIGURE 3-9 COMPARISON OF THE NON-LINEAR DEFLECTION AMONG THE THREE MICRO-FLEXURES

CONSIDERED CLAMPED-CLAMPED CRAB-LEG AND FOLDED BEAM ALL FLEXURES WERE DESIGNED TO HAVE THE SAME STIFFNESS OF 035 N M-1 26

FIGURE 3-10 LUMPED MASS APPROXIMATION WITH BASE EXCITATION YB(T) 27 FIGURE 3-11 LUMPED MASS APPROXIMATION WITH DIRECT MASS EXCITATION Z(T) 28 FIGURE 3-12 GRAPH REPRESENTING STABILITY REGIONS FOR GAP CLOSING CAPACITIVE DRIVES lsquorsquo

REPRESENTS THE RHS OF EQ (32) FOR T2 WHILE lsquorsquo REPRESENTS THE RHS OF T1 34 FIGURE 4-1 IDEAL CHARGE AMPLIFIER 36 FIGURE 4-2 CAPACITANCE TO VOLTAGE READOUT CIRCUIT WHICH IS MADE UP OF FOUR ELEMENTS THE

CHARGE AMPLIFIER VOLTAGE AMPLIFIER DEMODULATOR AND LOW PASS FILTER 37 FIGURE 4-3 ORCAD SIMULATION RESULTS OF THE CAPACITANCE TO VOLTAGE READOUT CIRCUIT THE

EXCITATION VOLTAGE VE IS 75 VPK AND THE OUTPUT OF THE VOLTAGE AMPLIFIER VVA IS APPROXIMATELY 13125 VPK 38

FIGURE 4-4 ORCAD SIMULATION RESULTS SHOWING THE SINUSOIDAL OUTPUT OF THE DEMODULATOR VDM WITH A 984 VPK-PK AMPLITUDE AND THE DC OUTPUT OF THE LOW PASS FILTER VLP AT APPROXIMATELY 5 V 39

FIGURE 5-1 ORCAD SIMULATION OF THE NOISE VOLTAGE SPECTRAL DENSITY ENO AT THE OUTPUT OF THE VOLTAGE AMPLIFIER 42

FIGURE 5-2 TOTAL OUTPUT NOISE ENO AT THE VOLTAGE AMPLIFIER WHICH IS EXPRESSED AS ENO=( int ENO2)12

42 FIGURE 6-1 POLYMUMPS FABRICATION PROCESS FOR A SIMPLE CANTILEVER BEAM 44 FIGURE 6-2 SOIMUMPS FABRICATION PROCESS FOR A SIMPLE CANTILEVER BEAM 45 FIGURE 7-1 THE SOLID LINE REPRESENTS THE THEORETICAL CAPACITANCE VERSUS DISPLACEMENT

RELATION WHILE THE SOLID POINTS REPRESENT THE EXPERIMENTAL VALUES THE DASHED LINE REPRESENTS THE ASSUMED LINEAR RELATION FOR DISPLACEMENTS LESS THAN 1ΜM THE EXPERIMENTAL AVERAGE DRIVE SENSITIVITY IS APPROXIMATELY 0024 PF ΜM-1 FOR DISPLACEMENTS GREATER THAN 12 ΜM DRIVE INSTABILITY IS OBSERVED AND IS ATTRIBUTED TO THE PULL-IN EFFECT 47

FIGURE 7-2 lsquotimesrsquo REPRESENTS THE AMPLITUDE AND SOLID LINE REPRESENTS THE PHASE RESPONSE THE RESONANCE PEAK IS LOCATED AT 3600 HZ AND ACCOMPANIED BY A PHASE SHIFT OF 168deg 48

FIGURE 7-3 A SEM MICRO-GRAPH SHOWING THE SENSOR TI FABRICATED USING POLYMUMPS THE FOUR FOLDED BEAM FLEXURES HAD A LENGTH OF 650 ΜM AND TOTAL STIFFNESS OF 14 NM THE EFFECTIVE PROOF MASS WAS 29E-9 KG AND THE SYSTEM HAD A NATURAL FREQUENCY OF 3500 HZ 50

FIGURE 7-4 A SEM MICRO-GRAPH OF THE TRANSVERSE COMB DRIVE IMPLEMENTED IN T1 WITH 200 COMB FINGERS A GAP DISTANCE OF 2 ΜM OVERLAP LENGTH OF 45 ΜM AND THICKNESS OF 2 ΜM THE

vi

AVERAGE EXPERIMENTAL SENSITIVITY WAS FOUND TO BE 0024 PF ΜM-1 FOR DISPLACEMENTS OF LESS THAN 1 ΜM 50

FIGURE 7-5 THE SOLID LINE REPRESENTS THE THEORETICAL CAPACITANCE TO VOLTAGE RELATION EXPRESSED BY EQ (37) WITH A SENSITIVITY OF 141 V PF-1 THE SOLID POINTS REPRESENT THE EXPERIMENTAL DATA POINTS WHEN DISCRETE CAPACITORS 1-4 PF ARE USED IN THE READOUT CIRCUIT THE HORIZONTAL DASHED LINE REPRESENTS THE OUTPUT VOLTAGE OF THE MEMS SENSOR AT REST AND CORRESPONDS TO A CAPACITANCE OF 35 PF 52

FIGURE 7-6 OSCILLOSCOPE TRACE OF THE INPUT EXCITATION VOLTAGE VE asymp 75 VPK AND OUTPUT OF THE VOLTAGE AMPLIFIER VVA asymp 103 VPK READINGS WERE TAKEN WITH THE MEMS SENSOR AT REST 53

FIGURE 7-7 OSCILLOSCOPE TRACE OF THE DEMODULATOR OUTPUT VDM asymp 76 VPK-PK AND OUTPUT OF THE LOW PASS FILTER VLP asymp 36 V READINGS WERE TAKEN WITH THE MEMS SENSOR AT REST 53

FIGURE 7-8 lsquorsquo REPRESENTS THE EXPERIMENTAL DATA POINTS WITH THE DASHED LINE REPRESENTING THE FITTED CURVE SOLID LINE REPRESENTS THE THEORETICAL DATA POINTS ACCORDING TO EQ (10) 54

FIGURE 7-9 EXPERIMENT SETUP WHICH INCLUDES THE MEMS SENOR AND READOUT BOARD MOUNTED ON THE CLAMPED-FREE BEAM AND TEST EQUIPMENT 55

FIGURE 7-10 CLOSE-UP OF THE MEMS SENSOR AND READOUT BOARD MOUNTED ON THE TEST BEAM ALONG WITH THE COMMERCIAL PIEZOELECTRIC ACCELEROMETER ONLY THE CHARGE AMPLIFIER (CA) AND THE VOLTAGE AMPLIFIER (VA) ARE MOUNTED ON THE SAME PROTOTYPE BOARD AS THE MEMS SENSOR DUE TO SPACE CONSTRAINTS 55

FIGURE 7-11 FIGURE COMPARING THE RESPONSE OF A COMMERCIAL PIEZOELECTRIC ACCELEROMETER AND MEMS SENSOR (L1) TO A BEAM VIBRATION A) DYNAMIC RESPONSE OF THE PIEZOELECTRIC ACCELEROMETER REPRESENTING THE TRUE BEAM VIBRATION WITH A PEAK VOLTAGE 114 V CORRESPONDING TO AN APPLIED ACCELERATION OF 114 G B) DYNAMIC RESPONSE OF THE MEMS SENSOR WITH A PEAK VOLTAGE OF 006 V 56

FIGURE 7-12 SENSOR DYNAMIC RESPONSE TO THE VIBRATION OF THE CLAMPED-FREE BEAM AND REPRESENTS A DIRECT EXPORT FROM THE DYNAMIC SIGNAL ANALYZER (AGILENT 35670) TOP TRACE IS THE TIME DOMAIN SIGNAL OF THE LOW PASS FILTER OUTPUT WHERE VLPMAXasymp73MV WHICH REPRESENTS A MAXIMUM ACCELERATION OF APPROXIMATELY 15G THE BOTTOM TRACE REPRESENTS THE FREQUENCY SPECTRUM WITH A PEAK AT 23HZ WHICH CORRESPONDS TO THE BEAM NATURAL FREQUENCY 57

FIGURE 7-13 EXPERIMENTAL RESULTS SHOWING THE RESPONSE OF THE SENSOR TO VARIOUS ACCELERATIONS IN THE RANGE OF 0-20 G WHERE lsquorsquo REPRESENTS THE EXPERIMENTAL DATA POINTS AND THE SOLID LINE REPRESENTS THE FITTED CURVE THE EXPERIMENTAL SENSITIVITY WAS FOUND TO BE 4258 MV G-1 58

vii

NOMENCLATURE a applied acceleration A applied acceleration amplitude App parallel plate overlapping area Asldie-film effective area for slide-film damping c coefficient of damping Cf feedback capacitance Cpp capacitance of the parallel plate capacitor cslide-film coefficient of damping as a result of slide-film damping

csqueeze-film coefficient of damping as a result of squeeze-film damping CTotal total capacitance of a comb drive CTotalBottom Total capacitance of the bottom drive in a tri-plate comb drive CTotalL total capacitance of a lateral drive CTotalT total capacitance of a transverse drive CTotalTop Total capacitance of the top drive in a tri-plate comb drive d distance between the parallel plates E Youngrsquos Modulus of sensor material eno noise voltage spectral density Eno total output noise voltage EnoAMP total output amplifier noise voltage EnoR total output resistor noise voltage

f excitation frequency in Hz fn system natural frequency in Hz

Fact electrostatic actuation force

FN noise force Fx-Elect total electrostatic force for pull-in instability Itruss second moment of inertia of truss section on the folded beam flexure Ibeam second moment of inertia of beam section on the folded beam

flexure kair dielectric constant of air at atmospheric pressure kB Boltzmannrsquos constant kT total stiffness of parallel folded beam flexure kx stiffness of various flexures in x-direction ky stiffness of various flexures in y-direction L flexure beam length Lplate effective plate length

viii

Ls shin beam length Lt thigh beam length m effective proof mass N number of comb fingers on the movable comb

NB number of balls

Qtotal total quality factor

Qf energy lost to surrounding fluid

Qs energy lost through supports

Qm energy lost through sensor material S theoretical sensitivity of sensor SL sensitivity of a lateral comb drive SM theoretical mechanical sensitivity of the proof massflexure system SRO sensitivity of readout circuit ST sensitivity of a transverse comb drive STAVG average sensitivity of a transverse comb drive t time T absolute temperature tb flexure beam thickness tc out of plane comb finger thickness

Vact electrostatic actuation voltage VCA output amplitude of charge amplifier VDM output amplitide of demodulator VE input amplitude of sinusoidal excitation voltage VLP output of low pass filter VLPMAX maximum output voltage at low pass filter VVA output amplitude of voltage amplifier vy relative motion of proof mass with respect to base Vy relative motion amplitude w flexure beam width Wplate effective plate width ws shin beam width wt thigh beam width x0 initial comb finger gap

x01 small gap distance on top side of the tri-plate comb drive

x02 small gap distance on bottom side of the tri-plate comb drive

ix

x03 larger gap distance on the tri-plate comb drive

XN noise displacement y0 initial comb finger overlap length yB base displacement YB base displacement amplitude ym absolute proof mass displacement α parameter dependant on the WplateLplate ratio ∆C Change in capacitance from a tri-plate drive in a differential

arrangement ε permittivity of free space κ second moment of inertia ratio (shinthigh) ζ damping ratio microair absolute viscosity of air ω base excitation frequency in rads ωn sensor natural frequency in rads ωv electrostatic actuation voltage frequency in rads L1 Sensor fabricated using SoiMUMPS with a lateral comb drive T1 Sensor fabricated using PolyMUMPS with a transverse comb drive T2 Sensor fabricated using SoiMUMPS with a tri-plate comb drive

x

1 INTRODUCTION

11 RESEARCH PURPOSE

The modern automobile consists of many mechanical systems such as power seats

windshield wipers mirrors trunks and windows which are all susceptible to breakdown

Without any condition monitoring system the breakdown is usually catastrophic and

requires an expensive part replacement Real-time condition monitoring allows for early

detection of faults which could require a simple solution such as the application of a

lubricant to fix This prolongs the useful life of the component and prevents sudden and

unexpected failure Real-time condition monitoring can be accomplished by examining the

vibration signature of a mechanical system For example an automobile power window

consists of a DC motor and its associated bearings and couplings a gear reduction system

consisting of worm and spur gears and kinematic links Faults resulting in excessive

vibrations may be caused by coupling misalignment bearing failure or gear train failure

Coupling misalignments occur at the connection between the drive shaft and the driven

shaft and are typically due to imperfect manufacturing Bearing failure is caused by lack

of lubrication or moisture contamination causing rusting while gear train failure is caused

by misaligned broken cracked or chipped gear teeth [1] Each fault occurs at a

characteristic frequency and so the state of the mechanical system can be determined by

monitoring the amplitudes of the relevant frequencies Vibrations due to coupling

misalignments occur at harmonics of the shaft rotational speed Gear vibrations occur at

the gear turn speed or at sidebands of the gear mesh frequency [1] Ball bearing vibrations

1

2

may be caused by outer bearing race defects fOB inner bearing race defects fIB ball

defects fB and train defects fT which all occur at specific frequencies [12]

cos( )12

dBOB r

p

bNf fd

β = minus

cos( )12

dBIB r

p

bNf fd

β = +

2cos( )1

2p r d

Bd p

d f bfb d

β = minus

cos( )12

drT

p

bffd

β = minus

(1)

where NB is the number of balls fr is the rotation frequency bd is the ball diameter β is the

contact angle between the ball and races and dp is the ball pitch diameter By monitoring

the real time vibration signature of the mechanical system anomalies can be quickly

identified and fixed

12 MICROELECTROMECHANICAL SENSORS

Microelectromechanical systems (MEMS) inertial sensors provide a small footprint

with sensitivities that are either comparable or exceed any macro sensor along with the

capability of mass production and low unit cost These sensors utilize compliant micro-

flexures attached to a proof mass that displaces in response to an environmental

acceleration Many transduction mechanisms have been developed that convert the

displacement into a measurable electric signal and include thermal piezoresistive

piezoelectric optical and capacitive methods Most MEMS sensors are silicon based and

are fabricated using surface or bulk micromachining Surface micromachining creates free

standing movable structures on top of a substrate using a combination of sacrificial layers

3

and structural layers which are commonly polysilicon [3] In bulk micromachining the

mechanical structures are defined using a removal process where bulk material typically

silicon is etched away MEMS sensors have been used for vibration and shock monitoring

on industrial systems and robotics guidance and navigation in global positioning systems

(GPS) seismometry in earthquake prediction image stabilization in digital cameras and

automobile safety and stability [4]

13 AUTOMOTIVE SENSORS

Sensors cover every major aspect of a modern automobile power-train sensors

monitor fuel combustion and emissions chassis sensors monitor road traction and tire

condition and body sensors facilitate air-bag deployment and vehicle proximity for radar

guided cruise control [5] Pressure sensors which typically consist of a piezoresistive

strain sensing element attached to a silicon diaphragm that deflects when exposed to an

applied pressure are one of the first micro-machined sensors used in an automobile

Implemented as manifold absolute pressure (MAP) sensors they allow precise control of

the air fuel ratio which allows the catalytic converter to efficiently reduce tailpipe

emissions [6] Variable reluctance sensors based on electromagnetics are used for

automobile traction control and produce a voltage output that is dependant on the

magnetic flux variations between a rotating component and the sensor bias magnet [57]

MEMS based linear accelerometers are utilized for airbag deployment upon impact and

provide a lower cost and smaller package over the older ball-in-tube accelerometers [6]

4

14 OPERATING ENVIRONMENT

Through proper packaging MEMS sensors can withstand the harsh automotive

environment where they face cyclic temperatures (-40degC to 150degC) mechanical shock or

vibration (10g-500g) and exposure to various fluids (oil brake fluid or engine coolant)

and airborne particles [8] Temperature changes induce thermal stresses causing

undesirable bending of the silicon microstructures that changes their dynamic response

Mechanical shock which could occur during a high impact collision or vibration from

uneven road conditions cause the delicate microstructures to fracture at points of stress

concentration These effects could be reduced by efficient design Use of a folded beam

micro-flexure that allows expansion and contraction or use of fillets at corners which

reduce the chance of fracture Bonding pads wire bonding and onboard circuit

components can be protected from air borne particles which could cause shorting using

parylene or gels Alternatively the entire package can be sealed hermetically which also

protects the MEMS device [8] In some cases such as fluid pressure sensors intimate

contact between the MEMS device and potentially corrosive fluid is necessary A recent

approach used a nanometer thin coating of silicon carbide (SiC) over a foundry fabricated

MEMS sensor using low temperature low pressure chemical vapor deposition (LPCVD)

[9] The SiC known for its chemical inertness did not compromise the mechanical

characteristics of the original design Therefore MEMS provides a strong platform for an

automotive sensor

5

15 OBJECTIVES AND CONTRIBUTION

The objective of this work is the development of a low frequency MEMS vibration

sensor The sensor must be low cost and mass producible which can be achieved using

standard foundry fabrication processes such as the Polysilicon Multi-User MEMS

Processes (PolyMUMPS) or the Silicon-on-Insulator Multi-User MEMS Processes

(SoiMUMPS) In addition the number of additional processing steps which are required

if integrated CMOS circuitry is used or if vacuum packaging is required must be

minimized This can be achieved using an off-chip circuit made with widely available

discrete components and sensor operation in air

In addition this work presents a MEMS sensor that is fabricated using standard

foundry processes such as PolyMUMPS and SoiMUMPS compared to many MEMS

sensors in literature that are fabricated using customized processes

16 RESEARCH OUTLINE

This work demonstrates three iterations of a low frequency capacitive vibration

sensor fabricated using the standard foundry processes Each sensor consists of a

suspended proof mass that displaces in response to an external vibration The proof mass

is attached to a folded beam micro-flexure compliant in one direction and its lateral

movement is sensed using a capacitive comb drive The first sensor T1 is fabricated in

PolyMUMPS and uses a transverse capacitive comb drive The second L1 fabricated in

SoiMUMPS uses a fully linear lateral capacitive drive and achieves an increase in

sensitivity from T1 Finally the third sensor T2 is fabricated in SoiMUMPS and uses a

differential tri-plate capacitive drive This sensor takes advantage of the increase in

6

sensitivity gained from a transverse drive but rejects off-axis displacement errors and

common mode using a differential tri-plate arrangement In addition an off-chip

capacitance to voltage readout circuit is fabricated and tested on sensor L1

Chapter 2 begins with an overview on the various methods of fault detection

followed by a review of MEMS vibration sensors each with a different transduction

mechanism Chapter 3 outlines the sensor structure focusing on the mechanical design

and sensing principal Chapter 4 focuses on the readout circuitry layout and simulation

results Chapter 5 outlines the various source of noise including mechanical noise arising

from the sensor and electric noise from the readout circuit Chapter 6 overviews the

foundry fabrication processes used including PolyMUMPs and SoiMUMPS The

experimental results of two sensors T1 and L1 which were fabricated and tested within

the thesis time frame are presented in Chapter 7 Chapter 8 concludes the thesis with a

brief summary and outline for future work

2 LITERATURE REVIEW

21 EXISTING METHODS FOR FAULT DETECTION

As this thesis develops a sensor for fault detection using vibration monitoring this

section overviews three existing methods for fault detection in mechanical systems First

motor current monitoring is discussed as it has the ability to measure significant bearing

faults in motor systems Next temperature sensors specifically thin film thermocouples

(TFTC) and MEMS temperature sensors are discussed These sensors are mounted in the

chamber of a lubricating fluid or in close proximity to a moving component and monitor

temperature fluctuations caused by increased wear Finally acoustic emission (AE)

sensors are discussed which have the ability to detect ultrasonic emissions generated by

plastic deformation in the crystalline lattice of a mechanical component

211 MOTOR CURRENT MONITORING

Monitoring of stator current is used as a method of determining bearing related faults

that produce radial motion between rotor and stator This motion varies the air gap flux

density producing stator currents at predictable frequencies fbng that represent the

bearing faults [10]

bng e vf f m f= plusmn sdot (2)

where fe is the electrical supply frequency m=123hellip and fv is one of the characteristics

frequencies listed in the equations of (1)The bearing damage must be significant to

produce noticeable current fluctuations and it may be difficult to distinguish from current

fluctuations due to voltage imbalances machine saturation or statorarmature faults [2]

7

8

212 TEMPERATURE SENSORS

Temperature sensors are largely used in high temperature high load and high speed

environments typically turbine engines and industrial tools Thermocouple sensors

consist of two dissimilar metals joined together at two junctions and when exposed to a

temperature difference generate a thermoelectric voltage dependant only on the material

properties and junction temperature difference [11] In turbine engines these sensors can

be placed on bearing casings or in the lubricant to monitor temperature fluctuations due

to increased friction The thermocouples are fabricated using bulk materials or micro-

machined thin films (TFTC) with the latter providing a faster response small package

and substantially lower production costs [12] The TFTC can be embedded for in situ

monitoring of processes in hostile environments One study successfully embedded

TFTC into electroplated nickel and attained device characteristics on par with surface

mounted devices [13] More recently MEMS temperature sensors have been developed

that consist of micro-machined semiconductor material that undergoes structural

deformation with temperature changes A common implementation is a bent beam

structure whose temperature induced deflection δ is expressed as [14]

( ) 3

2 2

sin12 cos sinz

EA T LEI AEL

α βδ 2β β

∆=

+

(3)

Temperature sensors are well suited for localized fault monitoring and detection but can

only characterize a mechanical system consisting of multiple components through the use

of multiple sensors

9

213 ACOUSTIC EMISSION SENSORS

Acoustic emission (AE) sensors have been used to characterize wear in machine

tools and monitor bearing and gear problems in centrifugal pumps [1516] First

developed as a Non-Destructive Testing (NDT) technique to detect cracks in civil

structures these sensors detect acoustic emissions generated by the release of vibration

waves in a crystalline lattice due to plastic deformation or crack growth [16]

Measurements are made using piezoelectric transducers with high natural frequencies

100 kHz to 1 MHz to capture the ultrasonic AE emissions An AE sensor is useful as it

has the ability to detect subsurface cracks in gear teeth or bearings before appearing on

the surface causing further damage [17] More recently MEMS acoustic sensors have

been developed and one design includes multiple transducers on a single substrate

which each detect acoustic emission energy at different frequencies [18] This helps

distinguish spurious acoustic emissions arising from impact and friction from those

arising from plastic deformation When compared to typical piezoelectric sensors the

MEMS devices have lower sensitivities and fail to detect some acoustic emissions

[1819] In addition the acoustic emission signal suffers severe attenuation as is crosses

various interfaces such as a gearbox or bearing casings In one experiment consisting of

a pinion gear and an associated bearing a 44dB attenuation was seen between an AE

sensor placed directly on the pinion to one placed on the bearing casing and in some

cases intermediate loss of the signal was observed [20]

10

22 MEMS VIBRATION SENSORS

This thesis focuses on the development of a MEMS vibration sensor for fault

detection as the vibration signature of a mechanical system has the potential to give an

overview of the entire system Many types of MEMS vibration sensors exist with

different operating principals based on resonance the piezoelectric effect and

displacement variation

221 RESONANT SENSORS

Resonant sensors are based on the resonant frequency shift of a fixed-fixed micro-

beam that is excited into resonance using electrostatic thermal or piezoelectric methods

Under an applied axial load there is a shift in the resonant frequency which is expressed

as [21]

2 2

2

1 12 1

nn n

EI Fl2

fl A Eα γ

π ρ= +

I

(4)

where l is the beam length E is Youngrsquos Modulus I is the moment of inertia ρ is the

density A is the cross section area F is the applied axial force αn and γn are mode

dependant coefficients and are 4730 and 0295 for mode 1 in the case discussed The

axial force F is generated by a proof mass attached to one beam end which moves in

response to an external acceleration The performance of resonant sensors are dependant

on the quality factor Q which is the ratio of total energy stored in the system to the

energy lost per cycle and determines the sharpness and amplitude of the resonance peak

A high Q resonator develops a peak that is easily distinguishable in a phase locked loop

control circuit and allows for improved performance and resolution [3] A resonator Q

value is lowered by energy losses to the surrounding fluid attached supports and

11

material Losses to the fluid surrounding the MEMS structure is the dominant loss

mechanism for MEMS sensors operated in air and can be greatly reduced by vacuum

packaging which can be accomplished using glass-silicon anodic bonding [22] In

addition to increasing fabrication cost the long term stability of the vacuum degrades

with ultimate device failure stemming from leakage through micro cracks and defects

and out-gassing [22] Support losses arise from the restoring forces they generate and

can be reduce by balanced designs such as a double ended tuning fork (DETF) or

operation in higher modes [3] In addition the sensitivity of a resonant sensor can be

increased by the addition of micro-levers to increase the axial force In one such design a

two stage micro-lever increased the sensitivity by an order of magnitude over a more

conventional single lever arm [2324]

222 PIEZOELECTRIC SENSORS

Piezoelectric materials produce a charge in response to an applied force This is an

intrinsic effect in materials such as Zinc Oxide (ZnO) whereas materials such as Barium

Titanate or Lead Zirconate Titanate (PZT) need to be poled by placing the material in a

strong electric field at an elevated temperature [3] The piezoelectric effect is expressed

mathematical by [25]

i ij j ik kD d T Eε= + (5)

where D represents the electrical displacement (Cm2) d is the charge coefficient (CN)

T is the applied stress matrix (Nm2) ε is the permittivity matrix (Fm) and E is the

electric field matrix (Vm) The subscripts i=123 j=123456 and k=123 represent

the direction to which the physical property is related For micro-sensors poled in the out

of plane direction only the D3 term is relevant and if fabricated using ZnO or PZT only

12

the charge coefficients d31 d32 and d33 are relevant [25] The piezoelectric material is

commonly deposited as a thin film on the surface of compliant structures in bulk micro-

machined sensors or incorporated as a layer in surface micro-machining on a similar

structure [26] However the inclusion of multiple materials introduces large temperature

variations causing out of plane bending especially in surface micro-machined devices

[2627] Poled piezoelectric materials could be depolarized which serves to reduce the

piezoelectric affect in the material This can occur if the sensor is exposed to a strong

electric field of the opposite polarity high temperatures in excess of the Curie point or

high mechanical stress Also the piezoelectric effect reduces over time an effect that can

be reduced with the addition of composite elements [28] Foundry processes such as

PolyMUMPS and SoiMUMPS do not incorporate piezoelectric layers thus sensors using

this method of transduction must use a customized process The piezoelectric sensor

outlined in [26] uses polysilicon surface micro-machining similar to PolyMUMPS

however it incorporates a ZnO layer that is not available in the foundry process

223 DISPLACEMENT VARIATION SENSORS

Displacement variation sensors consist of a proof mass connected to a compliant

micro-flexure The movement of the proof mass is sensed using optical electron

tunneling or more commonly capacitive methods A recent optical accelerometer

utilized a novel multilayer nano-grating for in-plane displacement sensing of the proof

mass [29] For small space variations between the nano-gratings a large change in the

optical reflection amplitude of the grating was observed [30] Even though optical

sensors offer resolutions approaching the Brownian noise limit they are not available in

small packages that are easily placed in space constricted areas for a low cost [29]

13

The electron tunneling effect is observed when the proximity between two electrodes

one flexible and the other fixed is on the order of 10Ǻ [31] If the two electrodes are

biased with voltage VB a tunneling current It flows between the gap [31]

exp( )t B I tgI V xαinfin sdot minus Φ (6)

where α1 is a constant and equal to 1025 Ǻ-1 eV-frac12 Ф is the effective height of the

tunneling barrier and xtg is the tunneling gap When the flexible electrode moves in

response to acceleration a feedback controller maintains the original distance while

determining the applied acceleration Electron tunneling sensors are able to sense

accelerations in the microg-ng range due to the exponential relation between tunneling current

and gap However these highly sensitive devices are susceptible to thermo-mechanical

noise which could be reduced by operation at low pressure [32] It is also important for

the two electrodes to be coated with metal which is difficult to achieve with foundry

fabrication

Capacitive sensors are based on the parallel plate capacitor and are implemented

using inter-digitated comb fingers in an in-plane lateral or transverse drive or in an out of

plane parallel plate drive [33] These sensors have the advantage of high sensitivity low

power consumption small package low temperature dependence and are easily

integrated with existing foundry fabrication processes [34] The lateral drive offers a fully

linear response but low sensitivity whereas the transverse and parallel plate drives

achieve high sensitivity if gap distances are of a few microns and large capacitance areas

are utilized However both drives have a non-linear response and are prone to pull-in

Also if these sensors are fabricated using a surface micromachining process such as

PolyMUMPS where the ground plane is a few microns above the structural layer they

14

are prone to electrical failure due to shorting This was observed in an experiment where

similar sensors were placed in a vibrating environment with a peak of 120g and shorting

between the structural layer and ground plane was found to be the predominate mode of

failure [35] However with certain fabrication processes such as SoiMUMPS this issue

can be neglected as the substrate material below any moving component is removed

Most capacitance sensors are fabricated using an integrated CMOS (Complementary

Metal Oxide Semiconductor)-MEMS technique where the MEMS sensor is fabricated

before during or after the CMOS circuit fabrication [36] Although these sensors offer

high-sensitivity and small parasitic capacitance they suffer from in-plane and out of

plane curling of the beam which reduces the capacitance between adjacent comb fingers

[37] In addition multiple processing steps are required for device fabrication

More than the other transduction methods MEMS capacitive displacement sensors are

attractive for practical implementation in an automobile and therefore is the selected

method for the vibration sensor presented in this work Their low temperature dependence

is an ideal characteristic for an automobile environment that faces varying temperature If

implemented with off-chip readout circuitry they require a standard foundry fabrication

process which provides a strong platform for mass production and low unit cost

3 SENSOR STRUCTURE

This chapter presents the physical structure of the three MEMS vibration sensors

which includes the displacement sensing capacitive comb drive and the micro-

flexureproof mass system First the capacitive comb drive and the three variations

implemented the transverse drive lateral drive and differential tri-plate drive are

discussed Next the mechanical structure is presented and includes the micro-flexure

selection quality factor and equations of motion Finally instability that places a limit

on the operating range of the sensor as a result of a gap closing capacitive comb drive is

discussed

31 SENSING

Capacitive sensors are based on the parallel plate capacitor and are commonly

implemented using inter-digitated comb fingers A capacitance can be realized if a voltage

is placed between two closely spaced plates If the fringing electric field is neglected the

capacitance of the parallel plate capacitor Cpp (F) is expressed as [28]

air pppp

k AC

dε

= (7)

where kair=1 is the dielectric constant of air at atmospheric pressure ε=88542times10-12 F m-1

is the permittivity of free space App is the overlapping area (m2) and d is the distance

between the two plates (m)

15

16

Figure 3-1 An inter-digitated comb drive where changes in capacitance are generated by either changes in gap distance x0 or in the overlap area y0timest

Figure 3-1

For a set of inter-digitated comb fingers or comb drive shown in the total

capacitance Ctotal is expressed as [28]

0 0

0 0

( ) ( )( ) (Total cy y y yC N t

)x x x xε

+ += + + minus

(8)

where N is the number of comb fingers on the movable comb y0 is the initial comb finger

overlap length (m) tc is the out of plane comb finger thickness (m) and x0 is the inital

finger gap (m) In this work the comb drive is implemented in three arrangements

transverse lateral and differential tri-plate In a transverse comb drive changes in

capacitance are generated by transverse movements along the x-axis resulting in changes

to the gap x0 between the movable and fixed comb fingers The nonlinear relation

between total capacitance CTotalT and displacement is expressed as

17

00 0

1 1( ) (Total TC N ty

)x x x xε

= + minus +

(9)

Although the drive sensitivity part is not constant an average sensitivity can be

assumed for small displacements [38]

Total TC xpart

In a lateral comb drive changes in capacitance are generated by lateral movements

along the y-axis resulting in changes to the overlapping area y0timestc The total capacitance

CTotalL between all comb fingers is expressed as

00

2 ( )Total LN tC yx

yε= +

(10)

The sensitivity of the lateral comb drive SL is expressed as

0

2Total LL

C N tSy x

εpart= =

part

(11)

The lateral comb drive was first introduced by W C Tang T H Nguyen and R T

Howe [33] in an electrostatic resonator and has since been used numerous times in both

drive and sense configurations It has the advantage of a fully linear response and

negligible instability but offers low sensitivity The sensitivity for the lateral drive L1

implemented in this work is 0038 pF microm-1

A differential tri-plate drive shown in Figure 3-2 which is a variation of the

transverse drive is implemented to improve sensitivity and reduce noise [39]

18

Figure 3-2 An inter-digitated comb drive in a differential tri-plate arrangement where x01 x02 and x03 are the gap distances and x01= x02ltlt x03 to improve sensitivity

In the tri-plate drive the at rest gap distances between the two capacitors formed by any

movable finger is not equivalent For a differential arrangement insuring that x01=

x02ltltx03 the linearity for small displacements (less than 1 microm) is enhanced over the case

of a single ended arrangement where x01=x02=x03=x0 which is equivalent to the

transverse drive This is shown graphically in Figure 3-3 where x01= x02=x03=2 microm in

the lower curve and x01= x02= 2 microm and x03= 15 microm in the upper curve A substantial

improvement in linearity is attained with the tri-plate drive

19

0

005

01

015

02

025

03

035

0 02 04 06 08 1

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-3 Comparison of the effect of gap distances on linearity for two case lsquorsquo represents case 1 where the fingers in are implemented in a differential tri-plate arrangement and x01= x02= 2 microm and x03= 15 microm lsquorsquo represents the case 2 where the fingers are implemented in a single ended arrangement and x01= x02=x03=2 microm

The tri-plate drive is also implemented in a differential arrangement where the output

is the change in capacitance rather than the total capacitance resulting from an external

excitation This reduces errors cause by off-axis excitation and common mode noise The

output of the differential tri-plate drive is expressed as [39]

Total Top Total BottomC C C∆ = minus where

003 01

1 1( ) (Total TopC N ty

)x x x xε

= +

minus +

(12)

20

002 03

1 1( ) (Total BottomC N ty

)x x x xε

= +

minus +

A graphic comparison of the capacitancedisplacement relation for sensors T1 and T2 is

shown in Figure 3-4 For these gap closing drives it is assumed that for displacements

less that 1 microm the relation is linear indicated by the dashed line above each curve The

sensitivity within this region is approximately 0046 pF microm-1 and 0253 pF microm-1 for T1

and T2 respectively The order of magnitude improvement is attributed to the differential

tri-plate arrangement and thicker comb finger

0

02

04

06

08

1

12

14

16

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-4 lsquorsquo represents the output of the differential tri-plate drive while lsquorsquo represents the output of the transverse drive The dashed line represents the linear approximation for displacements less that 1 microm For the purpose of comparison the output of the transverse drive was equalized to zero at zero displacement

21

In this work the transverse drive is first implemented in sensor T1 using the

PolyMUMPS fabrication process Then in an effort to linearize the sensor response the

lateral drive is used in L1 using SoiMUMPS The reduction in drive sensitivity due to the

lateral drive is further reduced by increasing the gap distance to 3 microm which is

recommended by the foundry to ensure a successful device These reductions are

balanced by increasing the number of comb fingers N and selecting a substrate thickness

that is 125 times large than T1 The third device T2 is fabricated using SoiMUMPS and

uses the tri-plate differential drive This design capitalizes on the increase in sensitivity of

a gap closing drive and larger substrate thickness attaining an order of magnitude

increase in sensitivity over T1 and L1 The specifications for each drive along with the

successive improvement in drive sensitivity are outlined in Table 3-1

Table 3-1 Specifications for each capacitive comb drive implemented in this work T1 L1 and T2

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Comb Finger Thickness tc (microm) 2 25 25 Gap Distance x0 (microm) 2 3 2 Number of Fingers N 200 256 30 Overlap length y0 (microm) 45 10 70 Drive Sensitivity (pF microm-1) 0046 0038 0273

32 MECHANICAL ANALYSIS

321 MICRO-FLEXURE SELECTION

The sensor structure consists of a proof mass connected to a compliant micro-flexure

that itself is anchored to the sensor substrate The proof mass displaces in response to an

22

applied acceleration and this displacement is sensed using the capacitive comb drive

attached to the proof mass

Proof mass

Folded beam micro-flexure

Anchor

Figure 3-5 Sensor Structure

This section focuses on micro-flexure selection as it should exhibit a large linear

displacement range and have a high lateral stiffness ratio to reduce cross-axis sensitivity

A number of micro-flexures have been used in MEMS research these include the

clamped-clamped crab-leg and folded beam shown in Figure 3-6-through 3-7

The clamped-clamped flexure exhibits a high stiffness ratio however for large lateral

displacements extensional axial forces buildup within the beams resulting in a non-linear

force-displacement relation [40] Within the linear regime the stiffness of the clamped-

clamped flexure is expressed as [40]

3

3b

yEt wk

L=

(13)

23

where E is the Youngrsquos Modulus of the beam material (N m-2) tb is the beam out of

plane thickness (m) w is the beam width (m) and L is the beam length (m) The stiffness

ratio is expressed as [40]

2x

y

k Lk w

=

(14)

If L is two orders of magnitude greater than w this results in a ratio of approximately

10000 The linear displacement range is determined using Finite Element Analysis (FEA)

of a clamped-clamped flexure with L=387 microm w=4 microm and tb=2 microm The result of this

simulation is plotted in and compared to a structure with constant stiffness ky

= 035 N m-1 predicted by Eq (13) The linear relation is valid for small deflections that

are less than 2 microm In addition this micro-flexure is susceptible to buckling as a result of

a compressive residual stress that may arise due to the fabrication process [41]

Figure 3-9

In the crab leg design the thigh section is meant to relieve the extensional axial

stress postponing the occurrence of non-linearity until greater lateral displacement [42]

The stiffness of the crab leg flexure is expressed as [43]

3

44

b s s ty

s s t

Et w L LkL L L

κκ

+= +

(15)

where κ=IsIt=(wswt)3 is a ratio of the second moment of inertia of the shin (s) and thigh

(t) ws and wt represent the width and Ls and Lt represent the length of the shin and thigh

regions respectively The stiffness ratio is expressed as [43]

34

4t s s tx

y s t s

w L L Lkk w L L L

αα t

+= +

(16)

The linear displacement range of a crab leg flexure with Ls=285 microm Lt=30 microm ws=wt=3

microm and tb=2 microm is determined using FEA The results are plotted in Figure 3-9 and

compared to a structure of constant stiffness ky=035 N m-1 predicted by Eq (15) The

24

non-linearity is greatly reduced when compared to the clamp-clamped flexure with Eq

accurately predicting the deflection until 4 microm However the flexure is more

susceptible to off-axis deflection as there is an order of magnitude reduction in the

stiffness ratio as compared to the clamped-clamped flexure

(15)

The folded beam flexure stands out as it is designed to exhibit a large linear

displacement range and to relieve residual stress due to the manufacturing process by

allowing the beams to expand and contract along the axial direction [40] It was first used

in an electrostatic resonator and has subsequently been used for electrostatic actuation in

a magnetic hard drives and micro-positioning [3344] If the truss section is rigid

ItrussgtgtIbeam then the stiffness of the flexure is expressed as [4043]

3

32b

yEt wk

L=

(17)

The stiffness ratio is expressed as [40]

2x

y

k Lk w

=

(18)

Figure 3-6 Clamped-clamped flexure

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

ACKNOWLEDGMENT

I would first like to thank God for all his blessings and express my gratitude to my

parents and brother for their endless support love and encouragement

I would like to thank my supervisors Dr William Cleghorn and Dr James Mills for

their guidance I would also like to thank my friends and colleagues Henry Chu Dr

Pezhman Hassanpour Shael Markin Dr Lidai Wang and Dr Xuping Zhang at the

Laboratory for Nonlinear Systems Control for great conversation guidance and

friendship Thank you to the members of my Examination Committee Dr Foued Ben

Amara and Dr Jean Zu for their time and feedback

Finally I would like to extend my appreciation to CMC Microsystems for the MEMS

device fabrication and design support and to Autorsquo21 a Network of Federal Centers of

Excellence for project funding

iii

TABLE OF CONTENTS 1 INTRODUCTION 1

11 RESEARCH PURPOSE 1 12 MICROELECTROMECHANICAL SENSORS 2 13 AUTOMOTIVE SENSORS 3 14 OPERATING ENVIRONMENT 4 15 OBJECTIVES AND CONTRIBUTION 5 16 RESEARCH OUTLINE 5

2 LITERATURE REVIEW 7

21 EXISTING METHODS FOR FAULT DETECTION 7 211 Motor Current Monitoring 7 212 Temperature Sensors 8 213 Acoustic Emission Sensors 9

22 MEMS VIBRATION SENSORS 10 221 Resonant Sensors 10 222 Piezoelectric Sensors 11 223 Displacement Variation Sensors 12

3 SENSOR STRUCTURE 15

31 SENSING 15 32 MECHANICAL ANALYSIS 21

321 Micro-Flexure Selection 21 322 Equations of Motion 26 323 Quality Factor 30

33 DRIVE STABILITY 32 34 ELECTROSTATIC ACTUATION 34 35 SECTION SUMMARY 35

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT 36

5 NOISE ANALYSIS 40

51 MECHANICAL-THERMAL NOISE 40 52 ELECTRICAL NOISE 41

6 FABRICATION OVERVIEW 43

61 POLYMUMPS 43 62 SOIMUMPS 44

7 EXPERIMENTAL RESULTS 46

71 MEMS SENSOR ndash T1 46 72 READOUT CIRCUITRY 51 73 MEMS SENSORndash L1 52 74 SECTION SUMMARY 59

8 CONCLUSION AND FUTURE WORK 61

81 CONCLUSION 61 82 FUTURE WORK 62

iv

LIST OF TABLES TABLE 3-1 SPECIFICATIONS FOR EACH CAPACITIVE COMB DRIVE IMPLEMENTED IN THIS WORK T1 L1 AND

T2 21 TABLE 3-2 DIMENSIONS AND MECHANICAL PROPERTIES FOR THE THREE MEMS VIBRATION SENSORS

OUTLINED IN THIS WORK 30 TABLE 3-3 SUMMARY OF THE THEORETICAL COEFFICIENTS OF DAMPING AND QUALITY FACTORS FOR THE

THREE SENSORS PRESENTED IN THIS WORK 32 TABLE 3-4 SUMMARY OF THE CHARACTERISTICS OF THE SENSORS OUTLINED IN THIS WORK 35 TABLE 7-1 SUMMARY OF EXPERIMENTAL RESULTS 60

v

LIST OF FIGURES

FIGURE 3-1 AN INTER-DIGITATED COMB DRIVE WHERE CHANGES IN CAPACITANCE ARE GENERATED BY

EITHER CHANGES IN GAP DISTANCE X0 OR IN THE OVERLAP AREA Y0timesT 16 FIGURE 3-2 AN INTER-DIGITATED COMB DRIVE IN A DIFFERENTIAL TRI-PLATE ARRANGEMENT WHERE X01

X02 AND X03 ARE THE GAP DISTANCES AND X01= X02ltlt X03 TO IMPROVE SENSITIVITY 18 FIGURE 3-3 COMPARISON OF THE EFFECT OF GAP DISTANCES ON LINEARITY FOR TWO CASE lsquorsquo REPRESENTS

CASE 1 WHERE THE FINGERS IN ARE IMPLEMENTED IN A DIFFERENTIAL TRI-PLATE ARRANGEMENT AND X01= X02= 2 ΜM AND X03= 15 ΜM lsquorsquo REPRESENTS THE CASE 2 WHERE THE FINGERS ARE IMPLEMENTED IN A SINGLE ENDED ARRANGEMENT AND X01= X02=X03=2 ΜM 19

FIGURE 3-4 lsquorsquo REPRESENTS THE OUTPUT OF THE DIFFERENTIAL TRI-PLATE DRIVE WHILE lsquorsquo REPRESENTS THE OUTPUT OF THE TRANSVERSE DRIVE THE DASHED LINE REPRESENTS THE LINEAR APPROXIMATION FOR DISPLACEMENTS LESS THAT 1 ΜM FOR THE PURPOSE OF COMPARISON THE OUTPUT OF THE TRANSVERSE DRIVE WAS EQUALIZED TO ZERO AT ZERO DISPLACEMENT 20

FIGURE 3-5 SENSOR STRUCTURE 22 FIGURE 3-6 CLAMPED-CLAMPED FLEXURE 24 FIGURE 3-7 CRAB LEG FLEXURE 25 FIGURE 3-8 FOLDED BEAM FLEXURE 25 FIGURE 3-9 COMPARISON OF THE NON-LINEAR DEFLECTION AMONG THE THREE MICRO-FLEXURES

CONSIDERED CLAMPED-CLAMPED CRAB-LEG AND FOLDED BEAM ALL FLEXURES WERE DESIGNED TO HAVE THE SAME STIFFNESS OF 035 N M-1 26

FIGURE 3-10 LUMPED MASS APPROXIMATION WITH BASE EXCITATION YB(T) 27 FIGURE 3-11 LUMPED MASS APPROXIMATION WITH DIRECT MASS EXCITATION Z(T) 28 FIGURE 3-12 GRAPH REPRESENTING STABILITY REGIONS FOR GAP CLOSING CAPACITIVE DRIVES lsquorsquo

REPRESENTS THE RHS OF EQ (32) FOR T2 WHILE lsquorsquo REPRESENTS THE RHS OF T1 34 FIGURE 4-1 IDEAL CHARGE AMPLIFIER 36 FIGURE 4-2 CAPACITANCE TO VOLTAGE READOUT CIRCUIT WHICH IS MADE UP OF FOUR ELEMENTS THE

CHARGE AMPLIFIER VOLTAGE AMPLIFIER DEMODULATOR AND LOW PASS FILTER 37 FIGURE 4-3 ORCAD SIMULATION RESULTS OF THE CAPACITANCE TO VOLTAGE READOUT CIRCUIT THE

EXCITATION VOLTAGE VE IS 75 VPK AND THE OUTPUT OF THE VOLTAGE AMPLIFIER VVA IS APPROXIMATELY 13125 VPK 38

FIGURE 4-4 ORCAD SIMULATION RESULTS SHOWING THE SINUSOIDAL OUTPUT OF THE DEMODULATOR VDM WITH A 984 VPK-PK AMPLITUDE AND THE DC OUTPUT OF THE LOW PASS FILTER VLP AT APPROXIMATELY 5 V 39

FIGURE 5-1 ORCAD SIMULATION OF THE NOISE VOLTAGE SPECTRAL DENSITY ENO AT THE OUTPUT OF THE VOLTAGE AMPLIFIER 42

FIGURE 5-2 TOTAL OUTPUT NOISE ENO AT THE VOLTAGE AMPLIFIER WHICH IS EXPRESSED AS ENO=( int ENO2)12

42 FIGURE 6-1 POLYMUMPS FABRICATION PROCESS FOR A SIMPLE CANTILEVER BEAM 44 FIGURE 6-2 SOIMUMPS FABRICATION PROCESS FOR A SIMPLE CANTILEVER BEAM 45 FIGURE 7-1 THE SOLID LINE REPRESENTS THE THEORETICAL CAPACITANCE VERSUS DISPLACEMENT

RELATION WHILE THE SOLID POINTS REPRESENT THE EXPERIMENTAL VALUES THE DASHED LINE REPRESENTS THE ASSUMED LINEAR RELATION FOR DISPLACEMENTS LESS THAN 1ΜM THE EXPERIMENTAL AVERAGE DRIVE SENSITIVITY IS APPROXIMATELY 0024 PF ΜM-1 FOR DISPLACEMENTS GREATER THAN 12 ΜM DRIVE INSTABILITY IS OBSERVED AND IS ATTRIBUTED TO THE PULL-IN EFFECT 47

FIGURE 7-2 lsquotimesrsquo REPRESENTS THE AMPLITUDE AND SOLID LINE REPRESENTS THE PHASE RESPONSE THE RESONANCE PEAK IS LOCATED AT 3600 HZ AND ACCOMPANIED BY A PHASE SHIFT OF 168deg 48

FIGURE 7-3 A SEM MICRO-GRAPH SHOWING THE SENSOR TI FABRICATED USING POLYMUMPS THE FOUR FOLDED BEAM FLEXURES HAD A LENGTH OF 650 ΜM AND TOTAL STIFFNESS OF 14 NM THE EFFECTIVE PROOF MASS WAS 29E-9 KG AND THE SYSTEM HAD A NATURAL FREQUENCY OF 3500 HZ 50

FIGURE 7-4 A SEM MICRO-GRAPH OF THE TRANSVERSE COMB DRIVE IMPLEMENTED IN T1 WITH 200 COMB FINGERS A GAP DISTANCE OF 2 ΜM OVERLAP LENGTH OF 45 ΜM AND THICKNESS OF 2 ΜM THE

vi

AVERAGE EXPERIMENTAL SENSITIVITY WAS FOUND TO BE 0024 PF ΜM-1 FOR DISPLACEMENTS OF LESS THAN 1 ΜM 50

FIGURE 7-5 THE SOLID LINE REPRESENTS THE THEORETICAL CAPACITANCE TO VOLTAGE RELATION EXPRESSED BY EQ (37) WITH A SENSITIVITY OF 141 V PF-1 THE SOLID POINTS REPRESENT THE EXPERIMENTAL DATA POINTS WHEN DISCRETE CAPACITORS 1-4 PF ARE USED IN THE READOUT CIRCUIT THE HORIZONTAL DASHED LINE REPRESENTS THE OUTPUT VOLTAGE OF THE MEMS SENSOR AT REST AND CORRESPONDS TO A CAPACITANCE OF 35 PF 52

FIGURE 7-6 OSCILLOSCOPE TRACE OF THE INPUT EXCITATION VOLTAGE VE asymp 75 VPK AND OUTPUT OF THE VOLTAGE AMPLIFIER VVA asymp 103 VPK READINGS WERE TAKEN WITH THE MEMS SENSOR AT REST 53

FIGURE 7-7 OSCILLOSCOPE TRACE OF THE DEMODULATOR OUTPUT VDM asymp 76 VPK-PK AND OUTPUT OF THE LOW PASS FILTER VLP asymp 36 V READINGS WERE TAKEN WITH THE MEMS SENSOR AT REST 53

FIGURE 7-8 lsquorsquo REPRESENTS THE EXPERIMENTAL DATA POINTS WITH THE DASHED LINE REPRESENTING THE FITTED CURVE SOLID LINE REPRESENTS THE THEORETICAL DATA POINTS ACCORDING TO EQ (10) 54

FIGURE 7-9 EXPERIMENT SETUP WHICH INCLUDES THE MEMS SENOR AND READOUT BOARD MOUNTED ON THE CLAMPED-FREE BEAM AND TEST EQUIPMENT 55

FIGURE 7-10 CLOSE-UP OF THE MEMS SENSOR AND READOUT BOARD MOUNTED ON THE TEST BEAM ALONG WITH THE COMMERCIAL PIEZOELECTRIC ACCELEROMETER ONLY THE CHARGE AMPLIFIER (CA) AND THE VOLTAGE AMPLIFIER (VA) ARE MOUNTED ON THE SAME PROTOTYPE BOARD AS THE MEMS SENSOR DUE TO SPACE CONSTRAINTS 55

FIGURE 7-11 FIGURE COMPARING THE RESPONSE OF A COMMERCIAL PIEZOELECTRIC ACCELEROMETER AND MEMS SENSOR (L1) TO A BEAM VIBRATION A) DYNAMIC RESPONSE OF THE PIEZOELECTRIC ACCELEROMETER REPRESENTING THE TRUE BEAM VIBRATION WITH A PEAK VOLTAGE 114 V CORRESPONDING TO AN APPLIED ACCELERATION OF 114 G B) DYNAMIC RESPONSE OF THE MEMS SENSOR WITH A PEAK VOLTAGE OF 006 V 56

FIGURE 7-12 SENSOR DYNAMIC RESPONSE TO THE VIBRATION OF THE CLAMPED-FREE BEAM AND REPRESENTS A DIRECT EXPORT FROM THE DYNAMIC SIGNAL ANALYZER (AGILENT 35670) TOP TRACE IS THE TIME DOMAIN SIGNAL OF THE LOW PASS FILTER OUTPUT WHERE VLPMAXasymp73MV WHICH REPRESENTS A MAXIMUM ACCELERATION OF APPROXIMATELY 15G THE BOTTOM TRACE REPRESENTS THE FREQUENCY SPECTRUM WITH A PEAK AT 23HZ WHICH CORRESPONDS TO THE BEAM NATURAL FREQUENCY 57

FIGURE 7-13 EXPERIMENTAL RESULTS SHOWING THE RESPONSE OF THE SENSOR TO VARIOUS ACCELERATIONS IN THE RANGE OF 0-20 G WHERE lsquorsquo REPRESENTS THE EXPERIMENTAL DATA POINTS AND THE SOLID LINE REPRESENTS THE FITTED CURVE THE EXPERIMENTAL SENSITIVITY WAS FOUND TO BE 4258 MV G-1 58

vii

NOMENCLATURE a applied acceleration A applied acceleration amplitude App parallel plate overlapping area Asldie-film effective area for slide-film damping c coefficient of damping Cf feedback capacitance Cpp capacitance of the parallel plate capacitor cslide-film coefficient of damping as a result of slide-film damping

csqueeze-film coefficient of damping as a result of squeeze-film damping CTotal total capacitance of a comb drive CTotalBottom Total capacitance of the bottom drive in a tri-plate comb drive CTotalL total capacitance of a lateral drive CTotalT total capacitance of a transverse drive CTotalTop Total capacitance of the top drive in a tri-plate comb drive d distance between the parallel plates E Youngrsquos Modulus of sensor material eno noise voltage spectral density Eno total output noise voltage EnoAMP total output amplifier noise voltage EnoR total output resistor noise voltage

f excitation frequency in Hz fn system natural frequency in Hz

Fact electrostatic actuation force

FN noise force Fx-Elect total electrostatic force for pull-in instability Itruss second moment of inertia of truss section on the folded beam flexure Ibeam second moment of inertia of beam section on the folded beam

flexure kair dielectric constant of air at atmospheric pressure kB Boltzmannrsquos constant kT total stiffness of parallel folded beam flexure kx stiffness of various flexures in x-direction ky stiffness of various flexures in y-direction L flexure beam length Lplate effective plate length

viii

Ls shin beam length Lt thigh beam length m effective proof mass N number of comb fingers on the movable comb

NB number of balls

Qtotal total quality factor

Qf energy lost to surrounding fluid

Qs energy lost through supports

Qm energy lost through sensor material S theoretical sensitivity of sensor SL sensitivity of a lateral comb drive SM theoretical mechanical sensitivity of the proof massflexure system SRO sensitivity of readout circuit ST sensitivity of a transverse comb drive STAVG average sensitivity of a transverse comb drive t time T absolute temperature tb flexure beam thickness tc out of plane comb finger thickness

Vact electrostatic actuation voltage VCA output amplitude of charge amplifier VDM output amplitide of demodulator VE input amplitude of sinusoidal excitation voltage VLP output of low pass filter VLPMAX maximum output voltage at low pass filter VVA output amplitude of voltage amplifier vy relative motion of proof mass with respect to base Vy relative motion amplitude w flexure beam width Wplate effective plate width ws shin beam width wt thigh beam width x0 initial comb finger gap

x01 small gap distance on top side of the tri-plate comb drive

x02 small gap distance on bottom side of the tri-plate comb drive

ix

x03 larger gap distance on the tri-plate comb drive

XN noise displacement y0 initial comb finger overlap length yB base displacement YB base displacement amplitude ym absolute proof mass displacement α parameter dependant on the WplateLplate ratio ∆C Change in capacitance from a tri-plate drive in a differential

arrangement ε permittivity of free space κ second moment of inertia ratio (shinthigh) ζ damping ratio microair absolute viscosity of air ω base excitation frequency in rads ωn sensor natural frequency in rads ωv electrostatic actuation voltage frequency in rads L1 Sensor fabricated using SoiMUMPS with a lateral comb drive T1 Sensor fabricated using PolyMUMPS with a transverse comb drive T2 Sensor fabricated using SoiMUMPS with a tri-plate comb drive

x

1 INTRODUCTION

11 RESEARCH PURPOSE

The modern automobile consists of many mechanical systems such as power seats

windshield wipers mirrors trunks and windows which are all susceptible to breakdown

Without any condition monitoring system the breakdown is usually catastrophic and

requires an expensive part replacement Real-time condition monitoring allows for early

detection of faults which could require a simple solution such as the application of a

lubricant to fix This prolongs the useful life of the component and prevents sudden and

unexpected failure Real-time condition monitoring can be accomplished by examining the

vibration signature of a mechanical system For example an automobile power window

consists of a DC motor and its associated bearings and couplings a gear reduction system

consisting of worm and spur gears and kinematic links Faults resulting in excessive

vibrations may be caused by coupling misalignment bearing failure or gear train failure

Coupling misalignments occur at the connection between the drive shaft and the driven

shaft and are typically due to imperfect manufacturing Bearing failure is caused by lack

of lubrication or moisture contamination causing rusting while gear train failure is caused

by misaligned broken cracked or chipped gear teeth [1] Each fault occurs at a

characteristic frequency and so the state of the mechanical system can be determined by

monitoring the amplitudes of the relevant frequencies Vibrations due to coupling

misalignments occur at harmonics of the shaft rotational speed Gear vibrations occur at

the gear turn speed or at sidebands of the gear mesh frequency [1] Ball bearing vibrations

1

2

may be caused by outer bearing race defects fOB inner bearing race defects fIB ball

defects fB and train defects fT which all occur at specific frequencies [12]

cos( )12

dBOB r

p

bNf fd

β = minus

cos( )12

dBIB r

p

bNf fd

β = +

2cos( )1

2p r d

Bd p

d f bfb d

β = minus

cos( )12

drT

p

bffd

β = minus

(1)

where NB is the number of balls fr is the rotation frequency bd is the ball diameter β is the

contact angle between the ball and races and dp is the ball pitch diameter By monitoring

the real time vibration signature of the mechanical system anomalies can be quickly

identified and fixed

12 MICROELECTROMECHANICAL SENSORS

Microelectromechanical systems (MEMS) inertial sensors provide a small footprint

with sensitivities that are either comparable or exceed any macro sensor along with the

capability of mass production and low unit cost These sensors utilize compliant micro-

flexures attached to a proof mass that displaces in response to an environmental

acceleration Many transduction mechanisms have been developed that convert the

displacement into a measurable electric signal and include thermal piezoresistive

piezoelectric optical and capacitive methods Most MEMS sensors are silicon based and

are fabricated using surface or bulk micromachining Surface micromachining creates free

standing movable structures on top of a substrate using a combination of sacrificial layers

3

and structural layers which are commonly polysilicon [3] In bulk micromachining the

mechanical structures are defined using a removal process where bulk material typically

silicon is etched away MEMS sensors have been used for vibration and shock monitoring

on industrial systems and robotics guidance and navigation in global positioning systems

(GPS) seismometry in earthquake prediction image stabilization in digital cameras and

automobile safety and stability [4]

13 AUTOMOTIVE SENSORS

Sensors cover every major aspect of a modern automobile power-train sensors

monitor fuel combustion and emissions chassis sensors monitor road traction and tire

condition and body sensors facilitate air-bag deployment and vehicle proximity for radar

guided cruise control [5] Pressure sensors which typically consist of a piezoresistive

strain sensing element attached to a silicon diaphragm that deflects when exposed to an

applied pressure are one of the first micro-machined sensors used in an automobile

Implemented as manifold absolute pressure (MAP) sensors they allow precise control of

the air fuel ratio which allows the catalytic converter to efficiently reduce tailpipe

emissions [6] Variable reluctance sensors based on electromagnetics are used for

automobile traction control and produce a voltage output that is dependant on the

magnetic flux variations between a rotating component and the sensor bias magnet [57]

MEMS based linear accelerometers are utilized for airbag deployment upon impact and

provide a lower cost and smaller package over the older ball-in-tube accelerometers [6]

4

14 OPERATING ENVIRONMENT

Through proper packaging MEMS sensors can withstand the harsh automotive

environment where they face cyclic temperatures (-40degC to 150degC) mechanical shock or

vibration (10g-500g) and exposure to various fluids (oil brake fluid or engine coolant)

and airborne particles [8] Temperature changes induce thermal stresses causing

undesirable bending of the silicon microstructures that changes their dynamic response

Mechanical shock which could occur during a high impact collision or vibration from

uneven road conditions cause the delicate microstructures to fracture at points of stress

concentration These effects could be reduced by efficient design Use of a folded beam

micro-flexure that allows expansion and contraction or use of fillets at corners which

reduce the chance of fracture Bonding pads wire bonding and onboard circuit

components can be protected from air borne particles which could cause shorting using

parylene or gels Alternatively the entire package can be sealed hermetically which also

protects the MEMS device [8] In some cases such as fluid pressure sensors intimate

contact between the MEMS device and potentially corrosive fluid is necessary A recent

approach used a nanometer thin coating of silicon carbide (SiC) over a foundry fabricated

MEMS sensor using low temperature low pressure chemical vapor deposition (LPCVD)

[9] The SiC known for its chemical inertness did not compromise the mechanical

characteristics of the original design Therefore MEMS provides a strong platform for an

automotive sensor

5

15 OBJECTIVES AND CONTRIBUTION

The objective of this work is the development of a low frequency MEMS vibration

sensor The sensor must be low cost and mass producible which can be achieved using

standard foundry fabrication processes such as the Polysilicon Multi-User MEMS

Processes (PolyMUMPS) or the Silicon-on-Insulator Multi-User MEMS Processes

(SoiMUMPS) In addition the number of additional processing steps which are required

if integrated CMOS circuitry is used or if vacuum packaging is required must be

minimized This can be achieved using an off-chip circuit made with widely available

discrete components and sensor operation in air

In addition this work presents a MEMS sensor that is fabricated using standard

foundry processes such as PolyMUMPS and SoiMUMPS compared to many MEMS

sensors in literature that are fabricated using customized processes

16 RESEARCH OUTLINE

This work demonstrates three iterations of a low frequency capacitive vibration

sensor fabricated using the standard foundry processes Each sensor consists of a

suspended proof mass that displaces in response to an external vibration The proof mass

is attached to a folded beam micro-flexure compliant in one direction and its lateral

movement is sensed using a capacitive comb drive The first sensor T1 is fabricated in

PolyMUMPS and uses a transverse capacitive comb drive The second L1 fabricated in

SoiMUMPS uses a fully linear lateral capacitive drive and achieves an increase in

sensitivity from T1 Finally the third sensor T2 is fabricated in SoiMUMPS and uses a

differential tri-plate capacitive drive This sensor takes advantage of the increase in

6

sensitivity gained from a transverse drive but rejects off-axis displacement errors and

common mode using a differential tri-plate arrangement In addition an off-chip

capacitance to voltage readout circuit is fabricated and tested on sensor L1

Chapter 2 begins with an overview on the various methods of fault detection

followed by a review of MEMS vibration sensors each with a different transduction

mechanism Chapter 3 outlines the sensor structure focusing on the mechanical design

and sensing principal Chapter 4 focuses on the readout circuitry layout and simulation

results Chapter 5 outlines the various source of noise including mechanical noise arising

from the sensor and electric noise from the readout circuit Chapter 6 overviews the

foundry fabrication processes used including PolyMUMPs and SoiMUMPS The

experimental results of two sensors T1 and L1 which were fabricated and tested within

the thesis time frame are presented in Chapter 7 Chapter 8 concludes the thesis with a

brief summary and outline for future work

2 LITERATURE REVIEW

21 EXISTING METHODS FOR FAULT DETECTION

As this thesis develops a sensor for fault detection using vibration monitoring this

section overviews three existing methods for fault detection in mechanical systems First

motor current monitoring is discussed as it has the ability to measure significant bearing

faults in motor systems Next temperature sensors specifically thin film thermocouples

(TFTC) and MEMS temperature sensors are discussed These sensors are mounted in the

chamber of a lubricating fluid or in close proximity to a moving component and monitor

temperature fluctuations caused by increased wear Finally acoustic emission (AE)

sensors are discussed which have the ability to detect ultrasonic emissions generated by

plastic deformation in the crystalline lattice of a mechanical component

211 MOTOR CURRENT MONITORING

Monitoring of stator current is used as a method of determining bearing related faults

that produce radial motion between rotor and stator This motion varies the air gap flux

density producing stator currents at predictable frequencies fbng that represent the

bearing faults [10]

bng e vf f m f= plusmn sdot (2)

where fe is the electrical supply frequency m=123hellip and fv is one of the characteristics

frequencies listed in the equations of (1)The bearing damage must be significant to

produce noticeable current fluctuations and it may be difficult to distinguish from current

fluctuations due to voltage imbalances machine saturation or statorarmature faults [2]

7

8

212 TEMPERATURE SENSORS

Temperature sensors are largely used in high temperature high load and high speed

environments typically turbine engines and industrial tools Thermocouple sensors

consist of two dissimilar metals joined together at two junctions and when exposed to a

temperature difference generate a thermoelectric voltage dependant only on the material

properties and junction temperature difference [11] In turbine engines these sensors can

be placed on bearing casings or in the lubricant to monitor temperature fluctuations due

to increased friction The thermocouples are fabricated using bulk materials or micro-

machined thin films (TFTC) with the latter providing a faster response small package

and substantially lower production costs [12] The TFTC can be embedded for in situ

monitoring of processes in hostile environments One study successfully embedded

TFTC into electroplated nickel and attained device characteristics on par with surface

mounted devices [13] More recently MEMS temperature sensors have been developed

that consist of micro-machined semiconductor material that undergoes structural

deformation with temperature changes A common implementation is a bent beam

structure whose temperature induced deflection δ is expressed as [14]

( ) 3

2 2

sin12 cos sinz

EA T LEI AEL

α βδ 2β β

∆=

+

(3)

Temperature sensors are well suited for localized fault monitoring and detection but can

only characterize a mechanical system consisting of multiple components through the use

of multiple sensors

9

213 ACOUSTIC EMISSION SENSORS

Acoustic emission (AE) sensors have been used to characterize wear in machine

tools and monitor bearing and gear problems in centrifugal pumps [1516] First

developed as a Non-Destructive Testing (NDT) technique to detect cracks in civil

structures these sensors detect acoustic emissions generated by the release of vibration

waves in a crystalline lattice due to plastic deformation or crack growth [16]

Measurements are made using piezoelectric transducers with high natural frequencies

100 kHz to 1 MHz to capture the ultrasonic AE emissions An AE sensor is useful as it

has the ability to detect subsurface cracks in gear teeth or bearings before appearing on

the surface causing further damage [17] More recently MEMS acoustic sensors have

been developed and one design includes multiple transducers on a single substrate

which each detect acoustic emission energy at different frequencies [18] This helps

distinguish spurious acoustic emissions arising from impact and friction from those

arising from plastic deformation When compared to typical piezoelectric sensors the

MEMS devices have lower sensitivities and fail to detect some acoustic emissions

[1819] In addition the acoustic emission signal suffers severe attenuation as is crosses

various interfaces such as a gearbox or bearing casings In one experiment consisting of

a pinion gear and an associated bearing a 44dB attenuation was seen between an AE

sensor placed directly on the pinion to one placed on the bearing casing and in some

cases intermediate loss of the signal was observed [20]

10

22 MEMS VIBRATION SENSORS

This thesis focuses on the development of a MEMS vibration sensor for fault

detection as the vibration signature of a mechanical system has the potential to give an

overview of the entire system Many types of MEMS vibration sensors exist with

different operating principals based on resonance the piezoelectric effect and

displacement variation

221 RESONANT SENSORS

Resonant sensors are based on the resonant frequency shift of a fixed-fixed micro-

beam that is excited into resonance using electrostatic thermal or piezoelectric methods

Under an applied axial load there is a shift in the resonant frequency which is expressed

as [21]

2 2

2

1 12 1

nn n

EI Fl2

fl A Eα γ

π ρ= +

I

(4)

where l is the beam length E is Youngrsquos Modulus I is the moment of inertia ρ is the

density A is the cross section area F is the applied axial force αn and γn are mode

dependant coefficients and are 4730 and 0295 for mode 1 in the case discussed The

axial force F is generated by a proof mass attached to one beam end which moves in

response to an external acceleration The performance of resonant sensors are dependant

on the quality factor Q which is the ratio of total energy stored in the system to the

energy lost per cycle and determines the sharpness and amplitude of the resonance peak

A high Q resonator develops a peak that is easily distinguishable in a phase locked loop

control circuit and allows for improved performance and resolution [3] A resonator Q

value is lowered by energy losses to the surrounding fluid attached supports and

11

material Losses to the fluid surrounding the MEMS structure is the dominant loss

mechanism for MEMS sensors operated in air and can be greatly reduced by vacuum

packaging which can be accomplished using glass-silicon anodic bonding [22] In

addition to increasing fabrication cost the long term stability of the vacuum degrades

with ultimate device failure stemming from leakage through micro cracks and defects

and out-gassing [22] Support losses arise from the restoring forces they generate and

can be reduce by balanced designs such as a double ended tuning fork (DETF) or

operation in higher modes [3] In addition the sensitivity of a resonant sensor can be

increased by the addition of micro-levers to increase the axial force In one such design a

two stage micro-lever increased the sensitivity by an order of magnitude over a more

conventional single lever arm [2324]

222 PIEZOELECTRIC SENSORS

Piezoelectric materials produce a charge in response to an applied force This is an

intrinsic effect in materials such as Zinc Oxide (ZnO) whereas materials such as Barium

Titanate or Lead Zirconate Titanate (PZT) need to be poled by placing the material in a

strong electric field at an elevated temperature [3] The piezoelectric effect is expressed

mathematical by [25]

i ij j ik kD d T Eε= + (5)

where D represents the electrical displacement (Cm2) d is the charge coefficient (CN)

T is the applied stress matrix (Nm2) ε is the permittivity matrix (Fm) and E is the

electric field matrix (Vm) The subscripts i=123 j=123456 and k=123 represent

the direction to which the physical property is related For micro-sensors poled in the out

of plane direction only the D3 term is relevant and if fabricated using ZnO or PZT only

12

the charge coefficients d31 d32 and d33 are relevant [25] The piezoelectric material is

commonly deposited as a thin film on the surface of compliant structures in bulk micro-

machined sensors or incorporated as a layer in surface micro-machining on a similar

structure [26] However the inclusion of multiple materials introduces large temperature

variations causing out of plane bending especially in surface micro-machined devices

[2627] Poled piezoelectric materials could be depolarized which serves to reduce the

piezoelectric affect in the material This can occur if the sensor is exposed to a strong

electric field of the opposite polarity high temperatures in excess of the Curie point or

high mechanical stress Also the piezoelectric effect reduces over time an effect that can

be reduced with the addition of composite elements [28] Foundry processes such as

PolyMUMPS and SoiMUMPS do not incorporate piezoelectric layers thus sensors using

this method of transduction must use a customized process The piezoelectric sensor

outlined in [26] uses polysilicon surface micro-machining similar to PolyMUMPS

however it incorporates a ZnO layer that is not available in the foundry process

223 DISPLACEMENT VARIATION SENSORS

Displacement variation sensors consist of a proof mass connected to a compliant

micro-flexure The movement of the proof mass is sensed using optical electron

tunneling or more commonly capacitive methods A recent optical accelerometer

utilized a novel multilayer nano-grating for in-plane displacement sensing of the proof

mass [29] For small space variations between the nano-gratings a large change in the

optical reflection amplitude of the grating was observed [30] Even though optical

sensors offer resolutions approaching the Brownian noise limit they are not available in

small packages that are easily placed in space constricted areas for a low cost [29]

13

The electron tunneling effect is observed when the proximity between two electrodes

one flexible and the other fixed is on the order of 10Ǻ [31] If the two electrodes are

biased with voltage VB a tunneling current It flows between the gap [31]

exp( )t B I tgI V xαinfin sdot minus Φ (6)

where α1 is a constant and equal to 1025 Ǻ-1 eV-frac12 Ф is the effective height of the

tunneling barrier and xtg is the tunneling gap When the flexible electrode moves in

response to acceleration a feedback controller maintains the original distance while

determining the applied acceleration Electron tunneling sensors are able to sense

accelerations in the microg-ng range due to the exponential relation between tunneling current

and gap However these highly sensitive devices are susceptible to thermo-mechanical

noise which could be reduced by operation at low pressure [32] It is also important for

the two electrodes to be coated with metal which is difficult to achieve with foundry

fabrication

Capacitive sensors are based on the parallel plate capacitor and are implemented

using inter-digitated comb fingers in an in-plane lateral or transverse drive or in an out of

plane parallel plate drive [33] These sensors have the advantage of high sensitivity low

power consumption small package low temperature dependence and are easily

integrated with existing foundry fabrication processes [34] The lateral drive offers a fully

linear response but low sensitivity whereas the transverse and parallel plate drives

achieve high sensitivity if gap distances are of a few microns and large capacitance areas

are utilized However both drives have a non-linear response and are prone to pull-in

Also if these sensors are fabricated using a surface micromachining process such as

PolyMUMPS where the ground plane is a few microns above the structural layer they

14

are prone to electrical failure due to shorting This was observed in an experiment where

similar sensors were placed in a vibrating environment with a peak of 120g and shorting

between the structural layer and ground plane was found to be the predominate mode of

failure [35] However with certain fabrication processes such as SoiMUMPS this issue

can be neglected as the substrate material below any moving component is removed

Most capacitance sensors are fabricated using an integrated CMOS (Complementary

Metal Oxide Semiconductor)-MEMS technique where the MEMS sensor is fabricated

before during or after the CMOS circuit fabrication [36] Although these sensors offer

high-sensitivity and small parasitic capacitance they suffer from in-plane and out of

plane curling of the beam which reduces the capacitance between adjacent comb fingers

[37] In addition multiple processing steps are required for device fabrication

More than the other transduction methods MEMS capacitive displacement sensors are

attractive for practical implementation in an automobile and therefore is the selected

method for the vibration sensor presented in this work Their low temperature dependence

is an ideal characteristic for an automobile environment that faces varying temperature If

implemented with off-chip readout circuitry they require a standard foundry fabrication

process which provides a strong platform for mass production and low unit cost

3 SENSOR STRUCTURE

This chapter presents the physical structure of the three MEMS vibration sensors

which includes the displacement sensing capacitive comb drive and the micro-

flexureproof mass system First the capacitive comb drive and the three variations

implemented the transverse drive lateral drive and differential tri-plate drive are

discussed Next the mechanical structure is presented and includes the micro-flexure

selection quality factor and equations of motion Finally instability that places a limit

on the operating range of the sensor as a result of a gap closing capacitive comb drive is

discussed

31 SENSING

Capacitive sensors are based on the parallel plate capacitor and are commonly

implemented using inter-digitated comb fingers A capacitance can be realized if a voltage

is placed between two closely spaced plates If the fringing electric field is neglected the

capacitance of the parallel plate capacitor Cpp (F) is expressed as [28]

air pppp

k AC

dε

= (7)

where kair=1 is the dielectric constant of air at atmospheric pressure ε=88542times10-12 F m-1

is the permittivity of free space App is the overlapping area (m2) and d is the distance

between the two plates (m)

15

16

Figure 3-1 An inter-digitated comb drive where changes in capacitance are generated by either changes in gap distance x0 or in the overlap area y0timest

Figure 3-1

For a set of inter-digitated comb fingers or comb drive shown in the total

capacitance Ctotal is expressed as [28]

0 0

0 0

( ) ( )( ) (Total cy y y yC N t

)x x x xε

+ += + + minus

(8)

where N is the number of comb fingers on the movable comb y0 is the initial comb finger

overlap length (m) tc is the out of plane comb finger thickness (m) and x0 is the inital

finger gap (m) In this work the comb drive is implemented in three arrangements

transverse lateral and differential tri-plate In a transverse comb drive changes in

capacitance are generated by transverse movements along the x-axis resulting in changes

to the gap x0 between the movable and fixed comb fingers The nonlinear relation

between total capacitance CTotalT and displacement is expressed as

17

00 0

1 1( ) (Total TC N ty

)x x x xε

= + minus +

(9)

Although the drive sensitivity part is not constant an average sensitivity can be

assumed for small displacements [38]

Total TC xpart

In a lateral comb drive changes in capacitance are generated by lateral movements

along the y-axis resulting in changes to the overlapping area y0timestc The total capacitance

CTotalL between all comb fingers is expressed as

00

2 ( )Total LN tC yx

yε= +

(10)

The sensitivity of the lateral comb drive SL is expressed as

0

2Total LL

C N tSy x

εpart= =

part

(11)

The lateral comb drive was first introduced by W C Tang T H Nguyen and R T

Howe [33] in an electrostatic resonator and has since been used numerous times in both

drive and sense configurations It has the advantage of a fully linear response and

negligible instability but offers low sensitivity The sensitivity for the lateral drive L1

implemented in this work is 0038 pF microm-1

A differential tri-plate drive shown in Figure 3-2 which is a variation of the

transverse drive is implemented to improve sensitivity and reduce noise [39]

18

Figure 3-2 An inter-digitated comb drive in a differential tri-plate arrangement where x01 x02 and x03 are the gap distances and x01= x02ltlt x03 to improve sensitivity

In the tri-plate drive the at rest gap distances between the two capacitors formed by any

movable finger is not equivalent For a differential arrangement insuring that x01=

x02ltltx03 the linearity for small displacements (less than 1 microm) is enhanced over the case

of a single ended arrangement where x01=x02=x03=x0 which is equivalent to the

transverse drive This is shown graphically in Figure 3-3 where x01= x02=x03=2 microm in

the lower curve and x01= x02= 2 microm and x03= 15 microm in the upper curve A substantial

improvement in linearity is attained with the tri-plate drive

19

0

005

01

015

02

025

03

035

0 02 04 06 08 1

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-3 Comparison of the effect of gap distances on linearity for two case lsquorsquo represents case 1 where the fingers in are implemented in a differential tri-plate arrangement and x01= x02= 2 microm and x03= 15 microm lsquorsquo represents the case 2 where the fingers are implemented in a single ended arrangement and x01= x02=x03=2 microm

The tri-plate drive is also implemented in a differential arrangement where the output

is the change in capacitance rather than the total capacitance resulting from an external

excitation This reduces errors cause by off-axis excitation and common mode noise The

output of the differential tri-plate drive is expressed as [39]

Total Top Total BottomC C C∆ = minus where

003 01

1 1( ) (Total TopC N ty

)x x x xε

= +

minus +

(12)

20

002 03

1 1( ) (Total BottomC N ty

)x x x xε

= +

minus +

A graphic comparison of the capacitancedisplacement relation for sensors T1 and T2 is

shown in Figure 3-4 For these gap closing drives it is assumed that for displacements

less that 1 microm the relation is linear indicated by the dashed line above each curve The

sensitivity within this region is approximately 0046 pF microm-1 and 0253 pF microm-1 for T1

and T2 respectively The order of magnitude improvement is attributed to the differential

tri-plate arrangement and thicker comb finger

0

02

04

06

08

1

12

14

16

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-4 lsquorsquo represents the output of the differential tri-plate drive while lsquorsquo represents the output of the transverse drive The dashed line represents the linear approximation for displacements less that 1 microm For the purpose of comparison the output of the transverse drive was equalized to zero at zero displacement

21

In this work the transverse drive is first implemented in sensor T1 using the

PolyMUMPS fabrication process Then in an effort to linearize the sensor response the

lateral drive is used in L1 using SoiMUMPS The reduction in drive sensitivity due to the

lateral drive is further reduced by increasing the gap distance to 3 microm which is

recommended by the foundry to ensure a successful device These reductions are

balanced by increasing the number of comb fingers N and selecting a substrate thickness

that is 125 times large than T1 The third device T2 is fabricated using SoiMUMPS and

uses the tri-plate differential drive This design capitalizes on the increase in sensitivity of

a gap closing drive and larger substrate thickness attaining an order of magnitude

increase in sensitivity over T1 and L1 The specifications for each drive along with the

successive improvement in drive sensitivity are outlined in Table 3-1

Table 3-1 Specifications for each capacitive comb drive implemented in this work T1 L1 and T2

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Comb Finger Thickness tc (microm) 2 25 25 Gap Distance x0 (microm) 2 3 2 Number of Fingers N 200 256 30 Overlap length y0 (microm) 45 10 70 Drive Sensitivity (pF microm-1) 0046 0038 0273

32 MECHANICAL ANALYSIS

321 MICRO-FLEXURE SELECTION

The sensor structure consists of a proof mass connected to a compliant micro-flexure

that itself is anchored to the sensor substrate The proof mass displaces in response to an

22

applied acceleration and this displacement is sensed using the capacitive comb drive

attached to the proof mass

Proof mass

Folded beam micro-flexure

Anchor

Figure 3-5 Sensor Structure

This section focuses on micro-flexure selection as it should exhibit a large linear

displacement range and have a high lateral stiffness ratio to reduce cross-axis sensitivity

A number of micro-flexures have been used in MEMS research these include the

clamped-clamped crab-leg and folded beam shown in Figure 3-6-through 3-7

The clamped-clamped flexure exhibits a high stiffness ratio however for large lateral

displacements extensional axial forces buildup within the beams resulting in a non-linear

force-displacement relation [40] Within the linear regime the stiffness of the clamped-

clamped flexure is expressed as [40]

3

3b

yEt wk

L=

(13)

23

where E is the Youngrsquos Modulus of the beam material (N m-2) tb is the beam out of

plane thickness (m) w is the beam width (m) and L is the beam length (m) The stiffness

ratio is expressed as [40]

2x

y

k Lk w

=

(14)

If L is two orders of magnitude greater than w this results in a ratio of approximately

10000 The linear displacement range is determined using Finite Element Analysis (FEA)

of a clamped-clamped flexure with L=387 microm w=4 microm and tb=2 microm The result of this

simulation is plotted in and compared to a structure with constant stiffness ky

= 035 N m-1 predicted by Eq (13) The linear relation is valid for small deflections that

are less than 2 microm In addition this micro-flexure is susceptible to buckling as a result of

a compressive residual stress that may arise due to the fabrication process [41]

Figure 3-9

In the crab leg design the thigh section is meant to relieve the extensional axial

stress postponing the occurrence of non-linearity until greater lateral displacement [42]

The stiffness of the crab leg flexure is expressed as [43]

3

44

b s s ty

s s t

Et w L LkL L L

κκ

+= +

(15)

where κ=IsIt=(wswt)3 is a ratio of the second moment of inertia of the shin (s) and thigh

(t) ws and wt represent the width and Ls and Lt represent the length of the shin and thigh

regions respectively The stiffness ratio is expressed as [43]

34

4t s s tx

y s t s

w L L Lkk w L L L

αα t

+= +

(16)

The linear displacement range of a crab leg flexure with Ls=285 microm Lt=30 microm ws=wt=3

microm and tb=2 microm is determined using FEA The results are plotted in Figure 3-9 and

compared to a structure of constant stiffness ky=035 N m-1 predicted by Eq (15) The

24

non-linearity is greatly reduced when compared to the clamp-clamped flexure with Eq

accurately predicting the deflection until 4 microm However the flexure is more

susceptible to off-axis deflection as there is an order of magnitude reduction in the

stiffness ratio as compared to the clamped-clamped flexure

(15)

The folded beam flexure stands out as it is designed to exhibit a large linear

displacement range and to relieve residual stress due to the manufacturing process by

allowing the beams to expand and contract along the axial direction [40] It was first used

in an electrostatic resonator and has subsequently been used for electrostatic actuation in

a magnetic hard drives and micro-positioning [3344] If the truss section is rigid

ItrussgtgtIbeam then the stiffness of the flexure is expressed as [4043]

3

32b

yEt wk

L=

(17)

The stiffness ratio is expressed as [40]

2x

y

k Lk w

=

(18)

Figure 3-6 Clamped-clamped flexure

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

TABLE OF CONTENTS 1 INTRODUCTION 1

11 RESEARCH PURPOSE 1 12 MICROELECTROMECHANICAL SENSORS 2 13 AUTOMOTIVE SENSORS 3 14 OPERATING ENVIRONMENT 4 15 OBJECTIVES AND CONTRIBUTION 5 16 RESEARCH OUTLINE 5

2 LITERATURE REVIEW 7

21 EXISTING METHODS FOR FAULT DETECTION 7 211 Motor Current Monitoring 7 212 Temperature Sensors 8 213 Acoustic Emission Sensors 9

22 MEMS VIBRATION SENSORS 10 221 Resonant Sensors 10 222 Piezoelectric Sensors 11 223 Displacement Variation Sensors 12

3 SENSOR STRUCTURE 15

31 SENSING 15 32 MECHANICAL ANALYSIS 21

321 Micro-Flexure Selection 21 322 Equations of Motion 26 323 Quality Factor 30

33 DRIVE STABILITY 32 34 ELECTROSTATIC ACTUATION 34 35 SECTION SUMMARY 35

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT 36

5 NOISE ANALYSIS 40

51 MECHANICAL-THERMAL NOISE 40 52 ELECTRICAL NOISE 41

6 FABRICATION OVERVIEW 43

61 POLYMUMPS 43 62 SOIMUMPS 44

7 EXPERIMENTAL RESULTS 46

71 MEMS SENSOR ndash T1 46 72 READOUT CIRCUITRY 51 73 MEMS SENSORndash L1 52 74 SECTION SUMMARY 59

8 CONCLUSION AND FUTURE WORK 61

81 CONCLUSION 61 82 FUTURE WORK 62

iv

LIST OF TABLES TABLE 3-1 SPECIFICATIONS FOR EACH CAPACITIVE COMB DRIVE IMPLEMENTED IN THIS WORK T1 L1 AND

T2 21 TABLE 3-2 DIMENSIONS AND MECHANICAL PROPERTIES FOR THE THREE MEMS VIBRATION SENSORS

OUTLINED IN THIS WORK 30 TABLE 3-3 SUMMARY OF THE THEORETICAL COEFFICIENTS OF DAMPING AND QUALITY FACTORS FOR THE

THREE SENSORS PRESENTED IN THIS WORK 32 TABLE 3-4 SUMMARY OF THE CHARACTERISTICS OF THE SENSORS OUTLINED IN THIS WORK 35 TABLE 7-1 SUMMARY OF EXPERIMENTAL RESULTS 60

v

LIST OF FIGURES

FIGURE 3-1 AN INTER-DIGITATED COMB DRIVE WHERE CHANGES IN CAPACITANCE ARE GENERATED BY

EITHER CHANGES IN GAP DISTANCE X0 OR IN THE OVERLAP AREA Y0timesT 16 FIGURE 3-2 AN INTER-DIGITATED COMB DRIVE IN A DIFFERENTIAL TRI-PLATE ARRANGEMENT WHERE X01

X02 AND X03 ARE THE GAP DISTANCES AND X01= X02ltlt X03 TO IMPROVE SENSITIVITY 18 FIGURE 3-3 COMPARISON OF THE EFFECT OF GAP DISTANCES ON LINEARITY FOR TWO CASE lsquorsquo REPRESENTS

CASE 1 WHERE THE FINGERS IN ARE IMPLEMENTED IN A DIFFERENTIAL TRI-PLATE ARRANGEMENT AND X01= X02= 2 ΜM AND X03= 15 ΜM lsquorsquo REPRESENTS THE CASE 2 WHERE THE FINGERS ARE IMPLEMENTED IN A SINGLE ENDED ARRANGEMENT AND X01= X02=X03=2 ΜM 19

FIGURE 3-4 lsquorsquo REPRESENTS THE OUTPUT OF THE DIFFERENTIAL TRI-PLATE DRIVE WHILE lsquorsquo REPRESENTS THE OUTPUT OF THE TRANSVERSE DRIVE THE DASHED LINE REPRESENTS THE LINEAR APPROXIMATION FOR DISPLACEMENTS LESS THAT 1 ΜM FOR THE PURPOSE OF COMPARISON THE OUTPUT OF THE TRANSVERSE DRIVE WAS EQUALIZED TO ZERO AT ZERO DISPLACEMENT 20

FIGURE 3-5 SENSOR STRUCTURE 22 FIGURE 3-6 CLAMPED-CLAMPED FLEXURE 24 FIGURE 3-7 CRAB LEG FLEXURE 25 FIGURE 3-8 FOLDED BEAM FLEXURE 25 FIGURE 3-9 COMPARISON OF THE NON-LINEAR DEFLECTION AMONG THE THREE MICRO-FLEXURES

CONSIDERED CLAMPED-CLAMPED CRAB-LEG AND FOLDED BEAM ALL FLEXURES WERE DESIGNED TO HAVE THE SAME STIFFNESS OF 035 N M-1 26

FIGURE 3-10 LUMPED MASS APPROXIMATION WITH BASE EXCITATION YB(T) 27 FIGURE 3-11 LUMPED MASS APPROXIMATION WITH DIRECT MASS EXCITATION Z(T) 28 FIGURE 3-12 GRAPH REPRESENTING STABILITY REGIONS FOR GAP CLOSING CAPACITIVE DRIVES lsquorsquo

REPRESENTS THE RHS OF EQ (32) FOR T2 WHILE lsquorsquo REPRESENTS THE RHS OF T1 34 FIGURE 4-1 IDEAL CHARGE AMPLIFIER 36 FIGURE 4-2 CAPACITANCE TO VOLTAGE READOUT CIRCUIT WHICH IS MADE UP OF FOUR ELEMENTS THE

CHARGE AMPLIFIER VOLTAGE AMPLIFIER DEMODULATOR AND LOW PASS FILTER 37 FIGURE 4-3 ORCAD SIMULATION RESULTS OF THE CAPACITANCE TO VOLTAGE READOUT CIRCUIT THE

EXCITATION VOLTAGE VE IS 75 VPK AND THE OUTPUT OF THE VOLTAGE AMPLIFIER VVA IS APPROXIMATELY 13125 VPK 38

FIGURE 4-4 ORCAD SIMULATION RESULTS SHOWING THE SINUSOIDAL OUTPUT OF THE DEMODULATOR VDM WITH A 984 VPK-PK AMPLITUDE AND THE DC OUTPUT OF THE LOW PASS FILTER VLP AT APPROXIMATELY 5 V 39

FIGURE 5-1 ORCAD SIMULATION OF THE NOISE VOLTAGE SPECTRAL DENSITY ENO AT THE OUTPUT OF THE VOLTAGE AMPLIFIER 42

FIGURE 5-2 TOTAL OUTPUT NOISE ENO AT THE VOLTAGE AMPLIFIER WHICH IS EXPRESSED AS ENO=( int ENO2)12

42 FIGURE 6-1 POLYMUMPS FABRICATION PROCESS FOR A SIMPLE CANTILEVER BEAM 44 FIGURE 6-2 SOIMUMPS FABRICATION PROCESS FOR A SIMPLE CANTILEVER BEAM 45 FIGURE 7-1 THE SOLID LINE REPRESENTS THE THEORETICAL CAPACITANCE VERSUS DISPLACEMENT

RELATION WHILE THE SOLID POINTS REPRESENT THE EXPERIMENTAL VALUES THE DASHED LINE REPRESENTS THE ASSUMED LINEAR RELATION FOR DISPLACEMENTS LESS THAN 1ΜM THE EXPERIMENTAL AVERAGE DRIVE SENSITIVITY IS APPROXIMATELY 0024 PF ΜM-1 FOR DISPLACEMENTS GREATER THAN 12 ΜM DRIVE INSTABILITY IS OBSERVED AND IS ATTRIBUTED TO THE PULL-IN EFFECT 47

FIGURE 7-2 lsquotimesrsquo REPRESENTS THE AMPLITUDE AND SOLID LINE REPRESENTS THE PHASE RESPONSE THE RESONANCE PEAK IS LOCATED AT 3600 HZ AND ACCOMPANIED BY A PHASE SHIFT OF 168deg 48

FIGURE 7-3 A SEM MICRO-GRAPH SHOWING THE SENSOR TI FABRICATED USING POLYMUMPS THE FOUR FOLDED BEAM FLEXURES HAD A LENGTH OF 650 ΜM AND TOTAL STIFFNESS OF 14 NM THE EFFECTIVE PROOF MASS WAS 29E-9 KG AND THE SYSTEM HAD A NATURAL FREQUENCY OF 3500 HZ 50

FIGURE 7-4 A SEM MICRO-GRAPH OF THE TRANSVERSE COMB DRIVE IMPLEMENTED IN T1 WITH 200 COMB FINGERS A GAP DISTANCE OF 2 ΜM OVERLAP LENGTH OF 45 ΜM AND THICKNESS OF 2 ΜM THE

vi

AVERAGE EXPERIMENTAL SENSITIVITY WAS FOUND TO BE 0024 PF ΜM-1 FOR DISPLACEMENTS OF LESS THAN 1 ΜM 50

FIGURE 7-5 THE SOLID LINE REPRESENTS THE THEORETICAL CAPACITANCE TO VOLTAGE RELATION EXPRESSED BY EQ (37) WITH A SENSITIVITY OF 141 V PF-1 THE SOLID POINTS REPRESENT THE EXPERIMENTAL DATA POINTS WHEN DISCRETE CAPACITORS 1-4 PF ARE USED IN THE READOUT CIRCUIT THE HORIZONTAL DASHED LINE REPRESENTS THE OUTPUT VOLTAGE OF THE MEMS SENSOR AT REST AND CORRESPONDS TO A CAPACITANCE OF 35 PF 52

FIGURE 7-6 OSCILLOSCOPE TRACE OF THE INPUT EXCITATION VOLTAGE VE asymp 75 VPK AND OUTPUT OF THE VOLTAGE AMPLIFIER VVA asymp 103 VPK READINGS WERE TAKEN WITH THE MEMS SENSOR AT REST 53

FIGURE 7-7 OSCILLOSCOPE TRACE OF THE DEMODULATOR OUTPUT VDM asymp 76 VPK-PK AND OUTPUT OF THE LOW PASS FILTER VLP asymp 36 V READINGS WERE TAKEN WITH THE MEMS SENSOR AT REST 53

FIGURE 7-8 lsquorsquo REPRESENTS THE EXPERIMENTAL DATA POINTS WITH THE DASHED LINE REPRESENTING THE FITTED CURVE SOLID LINE REPRESENTS THE THEORETICAL DATA POINTS ACCORDING TO EQ (10) 54

FIGURE 7-9 EXPERIMENT SETUP WHICH INCLUDES THE MEMS SENOR AND READOUT BOARD MOUNTED ON THE CLAMPED-FREE BEAM AND TEST EQUIPMENT 55

FIGURE 7-10 CLOSE-UP OF THE MEMS SENSOR AND READOUT BOARD MOUNTED ON THE TEST BEAM ALONG WITH THE COMMERCIAL PIEZOELECTRIC ACCELEROMETER ONLY THE CHARGE AMPLIFIER (CA) AND THE VOLTAGE AMPLIFIER (VA) ARE MOUNTED ON THE SAME PROTOTYPE BOARD AS THE MEMS SENSOR DUE TO SPACE CONSTRAINTS 55

FIGURE 7-11 FIGURE COMPARING THE RESPONSE OF A COMMERCIAL PIEZOELECTRIC ACCELEROMETER AND MEMS SENSOR (L1) TO A BEAM VIBRATION A) DYNAMIC RESPONSE OF THE PIEZOELECTRIC ACCELEROMETER REPRESENTING THE TRUE BEAM VIBRATION WITH A PEAK VOLTAGE 114 V CORRESPONDING TO AN APPLIED ACCELERATION OF 114 G B) DYNAMIC RESPONSE OF THE MEMS SENSOR WITH A PEAK VOLTAGE OF 006 V 56

FIGURE 7-12 SENSOR DYNAMIC RESPONSE TO THE VIBRATION OF THE CLAMPED-FREE BEAM AND REPRESENTS A DIRECT EXPORT FROM THE DYNAMIC SIGNAL ANALYZER (AGILENT 35670) TOP TRACE IS THE TIME DOMAIN SIGNAL OF THE LOW PASS FILTER OUTPUT WHERE VLPMAXasymp73MV WHICH REPRESENTS A MAXIMUM ACCELERATION OF APPROXIMATELY 15G THE BOTTOM TRACE REPRESENTS THE FREQUENCY SPECTRUM WITH A PEAK AT 23HZ WHICH CORRESPONDS TO THE BEAM NATURAL FREQUENCY 57

FIGURE 7-13 EXPERIMENTAL RESULTS SHOWING THE RESPONSE OF THE SENSOR TO VARIOUS ACCELERATIONS IN THE RANGE OF 0-20 G WHERE lsquorsquo REPRESENTS THE EXPERIMENTAL DATA POINTS AND THE SOLID LINE REPRESENTS THE FITTED CURVE THE EXPERIMENTAL SENSITIVITY WAS FOUND TO BE 4258 MV G-1 58

vii

NOMENCLATURE a applied acceleration A applied acceleration amplitude App parallel plate overlapping area Asldie-film effective area for slide-film damping c coefficient of damping Cf feedback capacitance Cpp capacitance of the parallel plate capacitor cslide-film coefficient of damping as a result of slide-film damping

csqueeze-film coefficient of damping as a result of squeeze-film damping CTotal total capacitance of a comb drive CTotalBottom Total capacitance of the bottom drive in a tri-plate comb drive CTotalL total capacitance of a lateral drive CTotalT total capacitance of a transverse drive CTotalTop Total capacitance of the top drive in a tri-plate comb drive d distance between the parallel plates E Youngrsquos Modulus of sensor material eno noise voltage spectral density Eno total output noise voltage EnoAMP total output amplifier noise voltage EnoR total output resistor noise voltage

f excitation frequency in Hz fn system natural frequency in Hz

Fact electrostatic actuation force

FN noise force Fx-Elect total electrostatic force for pull-in instability Itruss second moment of inertia of truss section on the folded beam flexure Ibeam second moment of inertia of beam section on the folded beam

flexure kair dielectric constant of air at atmospheric pressure kB Boltzmannrsquos constant kT total stiffness of parallel folded beam flexure kx stiffness of various flexures in x-direction ky stiffness of various flexures in y-direction L flexure beam length Lplate effective plate length

viii

Ls shin beam length Lt thigh beam length m effective proof mass N number of comb fingers on the movable comb

NB number of balls

Qtotal total quality factor

Qf energy lost to surrounding fluid

Qs energy lost through supports

Qm energy lost through sensor material S theoretical sensitivity of sensor SL sensitivity of a lateral comb drive SM theoretical mechanical sensitivity of the proof massflexure system SRO sensitivity of readout circuit ST sensitivity of a transverse comb drive STAVG average sensitivity of a transverse comb drive t time T absolute temperature tb flexure beam thickness tc out of plane comb finger thickness

Vact electrostatic actuation voltage VCA output amplitude of charge amplifier VDM output amplitide of demodulator VE input amplitude of sinusoidal excitation voltage VLP output of low pass filter VLPMAX maximum output voltage at low pass filter VVA output amplitude of voltage amplifier vy relative motion of proof mass with respect to base Vy relative motion amplitude w flexure beam width Wplate effective plate width ws shin beam width wt thigh beam width x0 initial comb finger gap

x01 small gap distance on top side of the tri-plate comb drive

x02 small gap distance on bottom side of the tri-plate comb drive

ix

x03 larger gap distance on the tri-plate comb drive

XN noise displacement y0 initial comb finger overlap length yB base displacement YB base displacement amplitude ym absolute proof mass displacement α parameter dependant on the WplateLplate ratio ∆C Change in capacitance from a tri-plate drive in a differential

arrangement ε permittivity of free space κ second moment of inertia ratio (shinthigh) ζ damping ratio microair absolute viscosity of air ω base excitation frequency in rads ωn sensor natural frequency in rads ωv electrostatic actuation voltage frequency in rads L1 Sensor fabricated using SoiMUMPS with a lateral comb drive T1 Sensor fabricated using PolyMUMPS with a transverse comb drive T2 Sensor fabricated using SoiMUMPS with a tri-plate comb drive

x

1 INTRODUCTION

11 RESEARCH PURPOSE

The modern automobile consists of many mechanical systems such as power seats

windshield wipers mirrors trunks and windows which are all susceptible to breakdown

Without any condition monitoring system the breakdown is usually catastrophic and

requires an expensive part replacement Real-time condition monitoring allows for early

detection of faults which could require a simple solution such as the application of a

lubricant to fix This prolongs the useful life of the component and prevents sudden and

unexpected failure Real-time condition monitoring can be accomplished by examining the

vibration signature of a mechanical system For example an automobile power window

consists of a DC motor and its associated bearings and couplings a gear reduction system

consisting of worm and spur gears and kinematic links Faults resulting in excessive

vibrations may be caused by coupling misalignment bearing failure or gear train failure

Coupling misalignments occur at the connection between the drive shaft and the driven

shaft and are typically due to imperfect manufacturing Bearing failure is caused by lack

of lubrication or moisture contamination causing rusting while gear train failure is caused

by misaligned broken cracked or chipped gear teeth [1] Each fault occurs at a

characteristic frequency and so the state of the mechanical system can be determined by

monitoring the amplitudes of the relevant frequencies Vibrations due to coupling

misalignments occur at harmonics of the shaft rotational speed Gear vibrations occur at

the gear turn speed or at sidebands of the gear mesh frequency [1] Ball bearing vibrations

1

2

may be caused by outer bearing race defects fOB inner bearing race defects fIB ball

defects fB and train defects fT which all occur at specific frequencies [12]

cos( )12

dBOB r

p

bNf fd

β = minus

cos( )12

dBIB r

p

bNf fd

β = +

2cos( )1

2p r d

Bd p

d f bfb d

β = minus

cos( )12

drT

p

bffd

β = minus

(1)

where NB is the number of balls fr is the rotation frequency bd is the ball diameter β is the

contact angle between the ball and races and dp is the ball pitch diameter By monitoring

the real time vibration signature of the mechanical system anomalies can be quickly

identified and fixed

12 MICROELECTROMECHANICAL SENSORS

Microelectromechanical systems (MEMS) inertial sensors provide a small footprint

with sensitivities that are either comparable or exceed any macro sensor along with the

capability of mass production and low unit cost These sensors utilize compliant micro-

flexures attached to a proof mass that displaces in response to an environmental

acceleration Many transduction mechanisms have been developed that convert the

displacement into a measurable electric signal and include thermal piezoresistive

piezoelectric optical and capacitive methods Most MEMS sensors are silicon based and

are fabricated using surface or bulk micromachining Surface micromachining creates free

standing movable structures on top of a substrate using a combination of sacrificial layers

3

and structural layers which are commonly polysilicon [3] In bulk micromachining the

mechanical structures are defined using a removal process where bulk material typically

silicon is etched away MEMS sensors have been used for vibration and shock monitoring

on industrial systems and robotics guidance and navigation in global positioning systems

(GPS) seismometry in earthquake prediction image stabilization in digital cameras and

automobile safety and stability [4]

13 AUTOMOTIVE SENSORS

Sensors cover every major aspect of a modern automobile power-train sensors

monitor fuel combustion and emissions chassis sensors monitor road traction and tire

condition and body sensors facilitate air-bag deployment and vehicle proximity for radar

guided cruise control [5] Pressure sensors which typically consist of a piezoresistive

strain sensing element attached to a silicon diaphragm that deflects when exposed to an

applied pressure are one of the first micro-machined sensors used in an automobile

Implemented as manifold absolute pressure (MAP) sensors they allow precise control of

the air fuel ratio which allows the catalytic converter to efficiently reduce tailpipe

emissions [6] Variable reluctance sensors based on electromagnetics are used for

automobile traction control and produce a voltage output that is dependant on the

magnetic flux variations between a rotating component and the sensor bias magnet [57]

MEMS based linear accelerometers are utilized for airbag deployment upon impact and

provide a lower cost and smaller package over the older ball-in-tube accelerometers [6]

4

14 OPERATING ENVIRONMENT

Through proper packaging MEMS sensors can withstand the harsh automotive

environment where they face cyclic temperatures (-40degC to 150degC) mechanical shock or

vibration (10g-500g) and exposure to various fluids (oil brake fluid or engine coolant)

and airborne particles [8] Temperature changes induce thermal stresses causing

undesirable bending of the silicon microstructures that changes their dynamic response

Mechanical shock which could occur during a high impact collision or vibration from

uneven road conditions cause the delicate microstructures to fracture at points of stress

concentration These effects could be reduced by efficient design Use of a folded beam

micro-flexure that allows expansion and contraction or use of fillets at corners which

reduce the chance of fracture Bonding pads wire bonding and onboard circuit

components can be protected from air borne particles which could cause shorting using

parylene or gels Alternatively the entire package can be sealed hermetically which also

protects the MEMS device [8] In some cases such as fluid pressure sensors intimate

contact between the MEMS device and potentially corrosive fluid is necessary A recent

approach used a nanometer thin coating of silicon carbide (SiC) over a foundry fabricated

MEMS sensor using low temperature low pressure chemical vapor deposition (LPCVD)

[9] The SiC known for its chemical inertness did not compromise the mechanical

characteristics of the original design Therefore MEMS provides a strong platform for an

automotive sensor

5

15 OBJECTIVES AND CONTRIBUTION

The objective of this work is the development of a low frequency MEMS vibration

sensor The sensor must be low cost and mass producible which can be achieved using

standard foundry fabrication processes such as the Polysilicon Multi-User MEMS

Processes (PolyMUMPS) or the Silicon-on-Insulator Multi-User MEMS Processes

(SoiMUMPS) In addition the number of additional processing steps which are required

if integrated CMOS circuitry is used or if vacuum packaging is required must be

minimized This can be achieved using an off-chip circuit made with widely available

discrete components and sensor operation in air

In addition this work presents a MEMS sensor that is fabricated using standard

foundry processes such as PolyMUMPS and SoiMUMPS compared to many MEMS

sensors in literature that are fabricated using customized processes

16 RESEARCH OUTLINE

This work demonstrates three iterations of a low frequency capacitive vibration

sensor fabricated using the standard foundry processes Each sensor consists of a

suspended proof mass that displaces in response to an external vibration The proof mass

is attached to a folded beam micro-flexure compliant in one direction and its lateral

movement is sensed using a capacitive comb drive The first sensor T1 is fabricated in

PolyMUMPS and uses a transverse capacitive comb drive The second L1 fabricated in

SoiMUMPS uses a fully linear lateral capacitive drive and achieves an increase in

sensitivity from T1 Finally the third sensor T2 is fabricated in SoiMUMPS and uses a

differential tri-plate capacitive drive This sensor takes advantage of the increase in

6

sensitivity gained from a transverse drive but rejects off-axis displacement errors and

common mode using a differential tri-plate arrangement In addition an off-chip

capacitance to voltage readout circuit is fabricated and tested on sensor L1

Chapter 2 begins with an overview on the various methods of fault detection

followed by a review of MEMS vibration sensors each with a different transduction

mechanism Chapter 3 outlines the sensor structure focusing on the mechanical design

and sensing principal Chapter 4 focuses on the readout circuitry layout and simulation

results Chapter 5 outlines the various source of noise including mechanical noise arising

from the sensor and electric noise from the readout circuit Chapter 6 overviews the

foundry fabrication processes used including PolyMUMPs and SoiMUMPS The

experimental results of two sensors T1 and L1 which were fabricated and tested within

the thesis time frame are presented in Chapter 7 Chapter 8 concludes the thesis with a

brief summary and outline for future work

2 LITERATURE REVIEW

21 EXISTING METHODS FOR FAULT DETECTION

As this thesis develops a sensor for fault detection using vibration monitoring this

section overviews three existing methods for fault detection in mechanical systems First

motor current monitoring is discussed as it has the ability to measure significant bearing

faults in motor systems Next temperature sensors specifically thin film thermocouples

(TFTC) and MEMS temperature sensors are discussed These sensors are mounted in the

chamber of a lubricating fluid or in close proximity to a moving component and monitor

temperature fluctuations caused by increased wear Finally acoustic emission (AE)

sensors are discussed which have the ability to detect ultrasonic emissions generated by

plastic deformation in the crystalline lattice of a mechanical component

211 MOTOR CURRENT MONITORING

Monitoring of stator current is used as a method of determining bearing related faults

that produce radial motion between rotor and stator This motion varies the air gap flux

density producing stator currents at predictable frequencies fbng that represent the

bearing faults [10]

bng e vf f m f= plusmn sdot (2)

where fe is the electrical supply frequency m=123hellip and fv is one of the characteristics

frequencies listed in the equations of (1)The bearing damage must be significant to

produce noticeable current fluctuations and it may be difficult to distinguish from current

fluctuations due to voltage imbalances machine saturation or statorarmature faults [2]

7

8

212 TEMPERATURE SENSORS

Temperature sensors are largely used in high temperature high load and high speed

environments typically turbine engines and industrial tools Thermocouple sensors

consist of two dissimilar metals joined together at two junctions and when exposed to a

temperature difference generate a thermoelectric voltage dependant only on the material

properties and junction temperature difference [11] In turbine engines these sensors can

be placed on bearing casings or in the lubricant to monitor temperature fluctuations due

to increased friction The thermocouples are fabricated using bulk materials or micro-

machined thin films (TFTC) with the latter providing a faster response small package

and substantially lower production costs [12] The TFTC can be embedded for in situ

monitoring of processes in hostile environments One study successfully embedded

TFTC into electroplated nickel and attained device characteristics on par with surface

mounted devices [13] More recently MEMS temperature sensors have been developed

that consist of micro-machined semiconductor material that undergoes structural

deformation with temperature changes A common implementation is a bent beam

structure whose temperature induced deflection δ is expressed as [14]

( ) 3

2 2

sin12 cos sinz

EA T LEI AEL

α βδ 2β β

∆=

+

(3)

Temperature sensors are well suited for localized fault monitoring and detection but can

only characterize a mechanical system consisting of multiple components through the use

of multiple sensors

9

213 ACOUSTIC EMISSION SENSORS

Acoustic emission (AE) sensors have been used to characterize wear in machine

tools and monitor bearing and gear problems in centrifugal pumps [1516] First

developed as a Non-Destructive Testing (NDT) technique to detect cracks in civil

structures these sensors detect acoustic emissions generated by the release of vibration

waves in a crystalline lattice due to plastic deformation or crack growth [16]

Measurements are made using piezoelectric transducers with high natural frequencies

100 kHz to 1 MHz to capture the ultrasonic AE emissions An AE sensor is useful as it

has the ability to detect subsurface cracks in gear teeth or bearings before appearing on

the surface causing further damage [17] More recently MEMS acoustic sensors have

been developed and one design includes multiple transducers on a single substrate

which each detect acoustic emission energy at different frequencies [18] This helps

distinguish spurious acoustic emissions arising from impact and friction from those

arising from plastic deformation When compared to typical piezoelectric sensors the

MEMS devices have lower sensitivities and fail to detect some acoustic emissions

[1819] In addition the acoustic emission signal suffers severe attenuation as is crosses

various interfaces such as a gearbox or bearing casings In one experiment consisting of

a pinion gear and an associated bearing a 44dB attenuation was seen between an AE

sensor placed directly on the pinion to one placed on the bearing casing and in some

cases intermediate loss of the signal was observed [20]

10

22 MEMS VIBRATION SENSORS

This thesis focuses on the development of a MEMS vibration sensor for fault

detection as the vibration signature of a mechanical system has the potential to give an

overview of the entire system Many types of MEMS vibration sensors exist with

different operating principals based on resonance the piezoelectric effect and

displacement variation

221 RESONANT SENSORS

Resonant sensors are based on the resonant frequency shift of a fixed-fixed micro-

beam that is excited into resonance using electrostatic thermal or piezoelectric methods

Under an applied axial load there is a shift in the resonant frequency which is expressed

as [21]

2 2

2

1 12 1

nn n

EI Fl2

fl A Eα γ

π ρ= +

I

(4)

where l is the beam length E is Youngrsquos Modulus I is the moment of inertia ρ is the

density A is the cross section area F is the applied axial force αn and γn are mode

dependant coefficients and are 4730 and 0295 for mode 1 in the case discussed The

axial force F is generated by a proof mass attached to one beam end which moves in

response to an external acceleration The performance of resonant sensors are dependant

on the quality factor Q which is the ratio of total energy stored in the system to the

energy lost per cycle and determines the sharpness and amplitude of the resonance peak

A high Q resonator develops a peak that is easily distinguishable in a phase locked loop

control circuit and allows for improved performance and resolution [3] A resonator Q

value is lowered by energy losses to the surrounding fluid attached supports and

11

material Losses to the fluid surrounding the MEMS structure is the dominant loss

mechanism for MEMS sensors operated in air and can be greatly reduced by vacuum

packaging which can be accomplished using glass-silicon anodic bonding [22] In

addition to increasing fabrication cost the long term stability of the vacuum degrades

with ultimate device failure stemming from leakage through micro cracks and defects

and out-gassing [22] Support losses arise from the restoring forces they generate and

can be reduce by balanced designs such as a double ended tuning fork (DETF) or

operation in higher modes [3] In addition the sensitivity of a resonant sensor can be

increased by the addition of micro-levers to increase the axial force In one such design a

two stage micro-lever increased the sensitivity by an order of magnitude over a more

conventional single lever arm [2324]

222 PIEZOELECTRIC SENSORS

Piezoelectric materials produce a charge in response to an applied force This is an

intrinsic effect in materials such as Zinc Oxide (ZnO) whereas materials such as Barium

Titanate or Lead Zirconate Titanate (PZT) need to be poled by placing the material in a

strong electric field at an elevated temperature [3] The piezoelectric effect is expressed

mathematical by [25]

i ij j ik kD d T Eε= + (5)

where D represents the electrical displacement (Cm2) d is the charge coefficient (CN)

T is the applied stress matrix (Nm2) ε is the permittivity matrix (Fm) and E is the

electric field matrix (Vm) The subscripts i=123 j=123456 and k=123 represent

the direction to which the physical property is related For micro-sensors poled in the out

of plane direction only the D3 term is relevant and if fabricated using ZnO or PZT only

12

the charge coefficients d31 d32 and d33 are relevant [25] The piezoelectric material is

commonly deposited as a thin film on the surface of compliant structures in bulk micro-

machined sensors or incorporated as a layer in surface micro-machining on a similar

structure [26] However the inclusion of multiple materials introduces large temperature

variations causing out of plane bending especially in surface micro-machined devices

[2627] Poled piezoelectric materials could be depolarized which serves to reduce the

piezoelectric affect in the material This can occur if the sensor is exposed to a strong

electric field of the opposite polarity high temperatures in excess of the Curie point or

high mechanical stress Also the piezoelectric effect reduces over time an effect that can

be reduced with the addition of composite elements [28] Foundry processes such as

PolyMUMPS and SoiMUMPS do not incorporate piezoelectric layers thus sensors using

this method of transduction must use a customized process The piezoelectric sensor

outlined in [26] uses polysilicon surface micro-machining similar to PolyMUMPS

however it incorporates a ZnO layer that is not available in the foundry process

223 DISPLACEMENT VARIATION SENSORS

Displacement variation sensors consist of a proof mass connected to a compliant

micro-flexure The movement of the proof mass is sensed using optical electron

tunneling or more commonly capacitive methods A recent optical accelerometer

utilized a novel multilayer nano-grating for in-plane displacement sensing of the proof

mass [29] For small space variations between the nano-gratings a large change in the

optical reflection amplitude of the grating was observed [30] Even though optical

sensors offer resolutions approaching the Brownian noise limit they are not available in

small packages that are easily placed in space constricted areas for a low cost [29]

13

The electron tunneling effect is observed when the proximity between two electrodes

one flexible and the other fixed is on the order of 10Ǻ [31] If the two electrodes are

biased with voltage VB a tunneling current It flows between the gap [31]

exp( )t B I tgI V xαinfin sdot minus Φ (6)

where α1 is a constant and equal to 1025 Ǻ-1 eV-frac12 Ф is the effective height of the

tunneling barrier and xtg is the tunneling gap When the flexible electrode moves in

response to acceleration a feedback controller maintains the original distance while

determining the applied acceleration Electron tunneling sensors are able to sense

accelerations in the microg-ng range due to the exponential relation between tunneling current

and gap However these highly sensitive devices are susceptible to thermo-mechanical

noise which could be reduced by operation at low pressure [32] It is also important for

the two electrodes to be coated with metal which is difficult to achieve with foundry

fabrication

Capacitive sensors are based on the parallel plate capacitor and are implemented

using inter-digitated comb fingers in an in-plane lateral or transverse drive or in an out of

plane parallel plate drive [33] These sensors have the advantage of high sensitivity low

power consumption small package low temperature dependence and are easily

integrated with existing foundry fabrication processes [34] The lateral drive offers a fully

linear response but low sensitivity whereas the transverse and parallel plate drives

achieve high sensitivity if gap distances are of a few microns and large capacitance areas

are utilized However both drives have a non-linear response and are prone to pull-in

Also if these sensors are fabricated using a surface micromachining process such as

PolyMUMPS where the ground plane is a few microns above the structural layer they

14

are prone to electrical failure due to shorting This was observed in an experiment where

similar sensors were placed in a vibrating environment with a peak of 120g and shorting

between the structural layer and ground plane was found to be the predominate mode of

failure [35] However with certain fabrication processes such as SoiMUMPS this issue

can be neglected as the substrate material below any moving component is removed

Most capacitance sensors are fabricated using an integrated CMOS (Complementary

Metal Oxide Semiconductor)-MEMS technique where the MEMS sensor is fabricated

before during or after the CMOS circuit fabrication [36] Although these sensors offer

high-sensitivity and small parasitic capacitance they suffer from in-plane and out of

plane curling of the beam which reduces the capacitance between adjacent comb fingers

[37] In addition multiple processing steps are required for device fabrication

More than the other transduction methods MEMS capacitive displacement sensors are

attractive for practical implementation in an automobile and therefore is the selected

method for the vibration sensor presented in this work Their low temperature dependence

is an ideal characteristic for an automobile environment that faces varying temperature If

implemented with off-chip readout circuitry they require a standard foundry fabrication

process which provides a strong platform for mass production and low unit cost

3 SENSOR STRUCTURE

This chapter presents the physical structure of the three MEMS vibration sensors

which includes the displacement sensing capacitive comb drive and the micro-

flexureproof mass system First the capacitive comb drive and the three variations

implemented the transverse drive lateral drive and differential tri-plate drive are

discussed Next the mechanical structure is presented and includes the micro-flexure

selection quality factor and equations of motion Finally instability that places a limit

on the operating range of the sensor as a result of a gap closing capacitive comb drive is

discussed

31 SENSING

Capacitive sensors are based on the parallel plate capacitor and are commonly

implemented using inter-digitated comb fingers A capacitance can be realized if a voltage

is placed between two closely spaced plates If the fringing electric field is neglected the

capacitance of the parallel plate capacitor Cpp (F) is expressed as [28]

air pppp

k AC

dε

= (7)

where kair=1 is the dielectric constant of air at atmospheric pressure ε=88542times10-12 F m-1

is the permittivity of free space App is the overlapping area (m2) and d is the distance

between the two plates (m)

15

16

Figure 3-1 An inter-digitated comb drive where changes in capacitance are generated by either changes in gap distance x0 or in the overlap area y0timest

Figure 3-1

For a set of inter-digitated comb fingers or comb drive shown in the total

capacitance Ctotal is expressed as [28]

0 0

0 0

( ) ( )( ) (Total cy y y yC N t

)x x x xε

+ += + + minus

(8)

where N is the number of comb fingers on the movable comb y0 is the initial comb finger

overlap length (m) tc is the out of plane comb finger thickness (m) and x0 is the inital

finger gap (m) In this work the comb drive is implemented in three arrangements

transverse lateral and differential tri-plate In a transverse comb drive changes in

capacitance are generated by transverse movements along the x-axis resulting in changes

to the gap x0 between the movable and fixed comb fingers The nonlinear relation

between total capacitance CTotalT and displacement is expressed as

17

00 0

1 1( ) (Total TC N ty

)x x x xε

= + minus +

(9)

Although the drive sensitivity part is not constant an average sensitivity can be

assumed for small displacements [38]

Total TC xpart

In a lateral comb drive changes in capacitance are generated by lateral movements

along the y-axis resulting in changes to the overlapping area y0timestc The total capacitance

CTotalL between all comb fingers is expressed as

00

2 ( )Total LN tC yx

yε= +

(10)

The sensitivity of the lateral comb drive SL is expressed as

0

2Total LL

C N tSy x

εpart= =

part

(11)

The lateral comb drive was first introduced by W C Tang T H Nguyen and R T

Howe [33] in an electrostatic resonator and has since been used numerous times in both

drive and sense configurations It has the advantage of a fully linear response and

negligible instability but offers low sensitivity The sensitivity for the lateral drive L1

implemented in this work is 0038 pF microm-1

A differential tri-plate drive shown in Figure 3-2 which is a variation of the

transverse drive is implemented to improve sensitivity and reduce noise [39]

18

Figure 3-2 An inter-digitated comb drive in a differential tri-plate arrangement where x01 x02 and x03 are the gap distances and x01= x02ltlt x03 to improve sensitivity

In the tri-plate drive the at rest gap distances between the two capacitors formed by any

movable finger is not equivalent For a differential arrangement insuring that x01=

x02ltltx03 the linearity for small displacements (less than 1 microm) is enhanced over the case

of a single ended arrangement where x01=x02=x03=x0 which is equivalent to the

transverse drive This is shown graphically in Figure 3-3 where x01= x02=x03=2 microm in

the lower curve and x01= x02= 2 microm and x03= 15 microm in the upper curve A substantial

improvement in linearity is attained with the tri-plate drive

19

0

005

01

015

02

025

03

035

0 02 04 06 08 1

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-3 Comparison of the effect of gap distances on linearity for two case lsquorsquo represents case 1 where the fingers in are implemented in a differential tri-plate arrangement and x01= x02= 2 microm and x03= 15 microm lsquorsquo represents the case 2 where the fingers are implemented in a single ended arrangement and x01= x02=x03=2 microm

The tri-plate drive is also implemented in a differential arrangement where the output

is the change in capacitance rather than the total capacitance resulting from an external

excitation This reduces errors cause by off-axis excitation and common mode noise The

output of the differential tri-plate drive is expressed as [39]

Total Top Total BottomC C C∆ = minus where

003 01

1 1( ) (Total TopC N ty

)x x x xε

= +

minus +

(12)

20

002 03

1 1( ) (Total BottomC N ty

)x x x xε

= +

minus +

A graphic comparison of the capacitancedisplacement relation for sensors T1 and T2 is

shown in Figure 3-4 For these gap closing drives it is assumed that for displacements

less that 1 microm the relation is linear indicated by the dashed line above each curve The

sensitivity within this region is approximately 0046 pF microm-1 and 0253 pF microm-1 for T1

and T2 respectively The order of magnitude improvement is attributed to the differential

tri-plate arrangement and thicker comb finger

0

02

04

06

08

1

12

14

16

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-4 lsquorsquo represents the output of the differential tri-plate drive while lsquorsquo represents the output of the transverse drive The dashed line represents the linear approximation for displacements less that 1 microm For the purpose of comparison the output of the transverse drive was equalized to zero at zero displacement

21

In this work the transverse drive is first implemented in sensor T1 using the

PolyMUMPS fabrication process Then in an effort to linearize the sensor response the

lateral drive is used in L1 using SoiMUMPS The reduction in drive sensitivity due to the

lateral drive is further reduced by increasing the gap distance to 3 microm which is

recommended by the foundry to ensure a successful device These reductions are

balanced by increasing the number of comb fingers N and selecting a substrate thickness

that is 125 times large than T1 The third device T2 is fabricated using SoiMUMPS and

uses the tri-plate differential drive This design capitalizes on the increase in sensitivity of

a gap closing drive and larger substrate thickness attaining an order of magnitude

increase in sensitivity over T1 and L1 The specifications for each drive along with the

successive improvement in drive sensitivity are outlined in Table 3-1

Table 3-1 Specifications for each capacitive comb drive implemented in this work T1 L1 and T2

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Comb Finger Thickness tc (microm) 2 25 25 Gap Distance x0 (microm) 2 3 2 Number of Fingers N 200 256 30 Overlap length y0 (microm) 45 10 70 Drive Sensitivity (pF microm-1) 0046 0038 0273

32 MECHANICAL ANALYSIS

321 MICRO-FLEXURE SELECTION

The sensor structure consists of a proof mass connected to a compliant micro-flexure

that itself is anchored to the sensor substrate The proof mass displaces in response to an

22

applied acceleration and this displacement is sensed using the capacitive comb drive

attached to the proof mass

Proof mass

Folded beam micro-flexure

Anchor

Figure 3-5 Sensor Structure

This section focuses on micro-flexure selection as it should exhibit a large linear

displacement range and have a high lateral stiffness ratio to reduce cross-axis sensitivity

A number of micro-flexures have been used in MEMS research these include the

clamped-clamped crab-leg and folded beam shown in Figure 3-6-through 3-7

The clamped-clamped flexure exhibits a high stiffness ratio however for large lateral

displacements extensional axial forces buildup within the beams resulting in a non-linear

force-displacement relation [40] Within the linear regime the stiffness of the clamped-

clamped flexure is expressed as [40]

3

3b

yEt wk

L=

(13)

23

where E is the Youngrsquos Modulus of the beam material (N m-2) tb is the beam out of

plane thickness (m) w is the beam width (m) and L is the beam length (m) The stiffness

ratio is expressed as [40]

2x

y

k Lk w

=

(14)

If L is two orders of magnitude greater than w this results in a ratio of approximately

10000 The linear displacement range is determined using Finite Element Analysis (FEA)

of a clamped-clamped flexure with L=387 microm w=4 microm and tb=2 microm The result of this

simulation is plotted in and compared to a structure with constant stiffness ky

= 035 N m-1 predicted by Eq (13) The linear relation is valid for small deflections that

are less than 2 microm In addition this micro-flexure is susceptible to buckling as a result of

a compressive residual stress that may arise due to the fabrication process [41]

Figure 3-9

In the crab leg design the thigh section is meant to relieve the extensional axial

stress postponing the occurrence of non-linearity until greater lateral displacement [42]

The stiffness of the crab leg flexure is expressed as [43]

3

44

b s s ty

s s t

Et w L LkL L L

κκ

+= +

(15)

where κ=IsIt=(wswt)3 is a ratio of the second moment of inertia of the shin (s) and thigh

(t) ws and wt represent the width and Ls and Lt represent the length of the shin and thigh

regions respectively The stiffness ratio is expressed as [43]

34

4t s s tx

y s t s

w L L Lkk w L L L

αα t

+= +

(16)

The linear displacement range of a crab leg flexure with Ls=285 microm Lt=30 microm ws=wt=3

microm and tb=2 microm is determined using FEA The results are plotted in Figure 3-9 and

compared to a structure of constant stiffness ky=035 N m-1 predicted by Eq (15) The

24

non-linearity is greatly reduced when compared to the clamp-clamped flexure with Eq

accurately predicting the deflection until 4 microm However the flexure is more

susceptible to off-axis deflection as there is an order of magnitude reduction in the

stiffness ratio as compared to the clamped-clamped flexure

(15)

The folded beam flexure stands out as it is designed to exhibit a large linear

displacement range and to relieve residual stress due to the manufacturing process by

allowing the beams to expand and contract along the axial direction [40] It was first used

in an electrostatic resonator and has subsequently been used for electrostatic actuation in

a magnetic hard drives and micro-positioning [3344] If the truss section is rigid

ItrussgtgtIbeam then the stiffness of the flexure is expressed as [4043]

3

32b

yEt wk

L=

(17)

The stiffness ratio is expressed as [40]

2x

y

k Lk w

=

(18)

Figure 3-6 Clamped-clamped flexure

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

LIST OF TABLES TABLE 3-1 SPECIFICATIONS FOR EACH CAPACITIVE COMB DRIVE IMPLEMENTED IN THIS WORK T1 L1 AND

T2 21 TABLE 3-2 DIMENSIONS AND MECHANICAL PROPERTIES FOR THE THREE MEMS VIBRATION SENSORS

OUTLINED IN THIS WORK 30 TABLE 3-3 SUMMARY OF THE THEORETICAL COEFFICIENTS OF DAMPING AND QUALITY FACTORS FOR THE

THREE SENSORS PRESENTED IN THIS WORK 32 TABLE 3-4 SUMMARY OF THE CHARACTERISTICS OF THE SENSORS OUTLINED IN THIS WORK 35 TABLE 7-1 SUMMARY OF EXPERIMENTAL RESULTS 60

v

LIST OF FIGURES

FIGURE 3-1 AN INTER-DIGITATED COMB DRIVE WHERE CHANGES IN CAPACITANCE ARE GENERATED BY

EITHER CHANGES IN GAP DISTANCE X0 OR IN THE OVERLAP AREA Y0timesT 16 FIGURE 3-2 AN INTER-DIGITATED COMB DRIVE IN A DIFFERENTIAL TRI-PLATE ARRANGEMENT WHERE X01

X02 AND X03 ARE THE GAP DISTANCES AND X01= X02ltlt X03 TO IMPROVE SENSITIVITY 18 FIGURE 3-3 COMPARISON OF THE EFFECT OF GAP DISTANCES ON LINEARITY FOR TWO CASE lsquorsquo REPRESENTS

CASE 1 WHERE THE FINGERS IN ARE IMPLEMENTED IN A DIFFERENTIAL TRI-PLATE ARRANGEMENT AND X01= X02= 2 ΜM AND X03= 15 ΜM lsquorsquo REPRESENTS THE CASE 2 WHERE THE FINGERS ARE IMPLEMENTED IN A SINGLE ENDED ARRANGEMENT AND X01= X02=X03=2 ΜM 19

FIGURE 3-4 lsquorsquo REPRESENTS THE OUTPUT OF THE DIFFERENTIAL TRI-PLATE DRIVE WHILE lsquorsquo REPRESENTS THE OUTPUT OF THE TRANSVERSE DRIVE THE DASHED LINE REPRESENTS THE LINEAR APPROXIMATION FOR DISPLACEMENTS LESS THAT 1 ΜM FOR THE PURPOSE OF COMPARISON THE OUTPUT OF THE TRANSVERSE DRIVE WAS EQUALIZED TO ZERO AT ZERO DISPLACEMENT 20

FIGURE 3-5 SENSOR STRUCTURE 22 FIGURE 3-6 CLAMPED-CLAMPED FLEXURE 24 FIGURE 3-7 CRAB LEG FLEXURE 25 FIGURE 3-8 FOLDED BEAM FLEXURE 25 FIGURE 3-9 COMPARISON OF THE NON-LINEAR DEFLECTION AMONG THE THREE MICRO-FLEXURES

CONSIDERED CLAMPED-CLAMPED CRAB-LEG AND FOLDED BEAM ALL FLEXURES WERE DESIGNED TO HAVE THE SAME STIFFNESS OF 035 N M-1 26

FIGURE 3-10 LUMPED MASS APPROXIMATION WITH BASE EXCITATION YB(T) 27 FIGURE 3-11 LUMPED MASS APPROXIMATION WITH DIRECT MASS EXCITATION Z(T) 28 FIGURE 3-12 GRAPH REPRESENTING STABILITY REGIONS FOR GAP CLOSING CAPACITIVE DRIVES lsquorsquo

REPRESENTS THE RHS OF EQ (32) FOR T2 WHILE lsquorsquo REPRESENTS THE RHS OF T1 34 FIGURE 4-1 IDEAL CHARGE AMPLIFIER 36 FIGURE 4-2 CAPACITANCE TO VOLTAGE READOUT CIRCUIT WHICH IS MADE UP OF FOUR ELEMENTS THE

CHARGE AMPLIFIER VOLTAGE AMPLIFIER DEMODULATOR AND LOW PASS FILTER 37 FIGURE 4-3 ORCAD SIMULATION RESULTS OF THE CAPACITANCE TO VOLTAGE READOUT CIRCUIT THE

EXCITATION VOLTAGE VE IS 75 VPK AND THE OUTPUT OF THE VOLTAGE AMPLIFIER VVA IS APPROXIMATELY 13125 VPK 38

FIGURE 4-4 ORCAD SIMULATION RESULTS SHOWING THE SINUSOIDAL OUTPUT OF THE DEMODULATOR VDM WITH A 984 VPK-PK AMPLITUDE AND THE DC OUTPUT OF THE LOW PASS FILTER VLP AT APPROXIMATELY 5 V 39

FIGURE 5-1 ORCAD SIMULATION OF THE NOISE VOLTAGE SPECTRAL DENSITY ENO AT THE OUTPUT OF THE VOLTAGE AMPLIFIER 42

FIGURE 5-2 TOTAL OUTPUT NOISE ENO AT THE VOLTAGE AMPLIFIER WHICH IS EXPRESSED AS ENO=( int ENO2)12

42 FIGURE 6-1 POLYMUMPS FABRICATION PROCESS FOR A SIMPLE CANTILEVER BEAM 44 FIGURE 6-2 SOIMUMPS FABRICATION PROCESS FOR A SIMPLE CANTILEVER BEAM 45 FIGURE 7-1 THE SOLID LINE REPRESENTS THE THEORETICAL CAPACITANCE VERSUS DISPLACEMENT

RELATION WHILE THE SOLID POINTS REPRESENT THE EXPERIMENTAL VALUES THE DASHED LINE REPRESENTS THE ASSUMED LINEAR RELATION FOR DISPLACEMENTS LESS THAN 1ΜM THE EXPERIMENTAL AVERAGE DRIVE SENSITIVITY IS APPROXIMATELY 0024 PF ΜM-1 FOR DISPLACEMENTS GREATER THAN 12 ΜM DRIVE INSTABILITY IS OBSERVED AND IS ATTRIBUTED TO THE PULL-IN EFFECT 47

FIGURE 7-2 lsquotimesrsquo REPRESENTS THE AMPLITUDE AND SOLID LINE REPRESENTS THE PHASE RESPONSE THE RESONANCE PEAK IS LOCATED AT 3600 HZ AND ACCOMPANIED BY A PHASE SHIFT OF 168deg 48

FIGURE 7-3 A SEM MICRO-GRAPH SHOWING THE SENSOR TI FABRICATED USING POLYMUMPS THE FOUR FOLDED BEAM FLEXURES HAD A LENGTH OF 650 ΜM AND TOTAL STIFFNESS OF 14 NM THE EFFECTIVE PROOF MASS WAS 29E-9 KG AND THE SYSTEM HAD A NATURAL FREQUENCY OF 3500 HZ 50

FIGURE 7-4 A SEM MICRO-GRAPH OF THE TRANSVERSE COMB DRIVE IMPLEMENTED IN T1 WITH 200 COMB FINGERS A GAP DISTANCE OF 2 ΜM OVERLAP LENGTH OF 45 ΜM AND THICKNESS OF 2 ΜM THE

vi

AVERAGE EXPERIMENTAL SENSITIVITY WAS FOUND TO BE 0024 PF ΜM-1 FOR DISPLACEMENTS OF LESS THAN 1 ΜM 50

FIGURE 7-5 THE SOLID LINE REPRESENTS THE THEORETICAL CAPACITANCE TO VOLTAGE RELATION EXPRESSED BY EQ (37) WITH A SENSITIVITY OF 141 V PF-1 THE SOLID POINTS REPRESENT THE EXPERIMENTAL DATA POINTS WHEN DISCRETE CAPACITORS 1-4 PF ARE USED IN THE READOUT CIRCUIT THE HORIZONTAL DASHED LINE REPRESENTS THE OUTPUT VOLTAGE OF THE MEMS SENSOR AT REST AND CORRESPONDS TO A CAPACITANCE OF 35 PF 52

FIGURE 7-6 OSCILLOSCOPE TRACE OF THE INPUT EXCITATION VOLTAGE VE asymp 75 VPK AND OUTPUT OF THE VOLTAGE AMPLIFIER VVA asymp 103 VPK READINGS WERE TAKEN WITH THE MEMS SENSOR AT REST 53

FIGURE 7-7 OSCILLOSCOPE TRACE OF THE DEMODULATOR OUTPUT VDM asymp 76 VPK-PK AND OUTPUT OF THE LOW PASS FILTER VLP asymp 36 V READINGS WERE TAKEN WITH THE MEMS SENSOR AT REST 53

FIGURE 7-8 lsquorsquo REPRESENTS THE EXPERIMENTAL DATA POINTS WITH THE DASHED LINE REPRESENTING THE FITTED CURVE SOLID LINE REPRESENTS THE THEORETICAL DATA POINTS ACCORDING TO EQ (10) 54

FIGURE 7-9 EXPERIMENT SETUP WHICH INCLUDES THE MEMS SENOR AND READOUT BOARD MOUNTED ON THE CLAMPED-FREE BEAM AND TEST EQUIPMENT 55

FIGURE 7-10 CLOSE-UP OF THE MEMS SENSOR AND READOUT BOARD MOUNTED ON THE TEST BEAM ALONG WITH THE COMMERCIAL PIEZOELECTRIC ACCELEROMETER ONLY THE CHARGE AMPLIFIER (CA) AND THE VOLTAGE AMPLIFIER (VA) ARE MOUNTED ON THE SAME PROTOTYPE BOARD AS THE MEMS SENSOR DUE TO SPACE CONSTRAINTS 55

FIGURE 7-11 FIGURE COMPARING THE RESPONSE OF A COMMERCIAL PIEZOELECTRIC ACCELEROMETER AND MEMS SENSOR (L1) TO A BEAM VIBRATION A) DYNAMIC RESPONSE OF THE PIEZOELECTRIC ACCELEROMETER REPRESENTING THE TRUE BEAM VIBRATION WITH A PEAK VOLTAGE 114 V CORRESPONDING TO AN APPLIED ACCELERATION OF 114 G B) DYNAMIC RESPONSE OF THE MEMS SENSOR WITH A PEAK VOLTAGE OF 006 V 56

FIGURE 7-12 SENSOR DYNAMIC RESPONSE TO THE VIBRATION OF THE CLAMPED-FREE BEAM AND REPRESENTS A DIRECT EXPORT FROM THE DYNAMIC SIGNAL ANALYZER (AGILENT 35670) TOP TRACE IS THE TIME DOMAIN SIGNAL OF THE LOW PASS FILTER OUTPUT WHERE VLPMAXasymp73MV WHICH REPRESENTS A MAXIMUM ACCELERATION OF APPROXIMATELY 15G THE BOTTOM TRACE REPRESENTS THE FREQUENCY SPECTRUM WITH A PEAK AT 23HZ WHICH CORRESPONDS TO THE BEAM NATURAL FREQUENCY 57

FIGURE 7-13 EXPERIMENTAL RESULTS SHOWING THE RESPONSE OF THE SENSOR TO VARIOUS ACCELERATIONS IN THE RANGE OF 0-20 G WHERE lsquorsquo REPRESENTS THE EXPERIMENTAL DATA POINTS AND THE SOLID LINE REPRESENTS THE FITTED CURVE THE EXPERIMENTAL SENSITIVITY WAS FOUND TO BE 4258 MV G-1 58

vii

NOMENCLATURE a applied acceleration A applied acceleration amplitude App parallel plate overlapping area Asldie-film effective area for slide-film damping c coefficient of damping Cf feedback capacitance Cpp capacitance of the parallel plate capacitor cslide-film coefficient of damping as a result of slide-film damping

csqueeze-film coefficient of damping as a result of squeeze-film damping CTotal total capacitance of a comb drive CTotalBottom Total capacitance of the bottom drive in a tri-plate comb drive CTotalL total capacitance of a lateral drive CTotalT total capacitance of a transverse drive CTotalTop Total capacitance of the top drive in a tri-plate comb drive d distance between the parallel plates E Youngrsquos Modulus of sensor material eno noise voltage spectral density Eno total output noise voltage EnoAMP total output amplifier noise voltage EnoR total output resistor noise voltage

f excitation frequency in Hz fn system natural frequency in Hz

Fact electrostatic actuation force

FN noise force Fx-Elect total electrostatic force for pull-in instability Itruss second moment of inertia of truss section on the folded beam flexure Ibeam second moment of inertia of beam section on the folded beam

flexure kair dielectric constant of air at atmospheric pressure kB Boltzmannrsquos constant kT total stiffness of parallel folded beam flexure kx stiffness of various flexures in x-direction ky stiffness of various flexures in y-direction L flexure beam length Lplate effective plate length

viii

Ls shin beam length Lt thigh beam length m effective proof mass N number of comb fingers on the movable comb

NB number of balls

Qtotal total quality factor

Qf energy lost to surrounding fluid

Qs energy lost through supports

Qm energy lost through sensor material S theoretical sensitivity of sensor SL sensitivity of a lateral comb drive SM theoretical mechanical sensitivity of the proof massflexure system SRO sensitivity of readout circuit ST sensitivity of a transverse comb drive STAVG average sensitivity of a transverse comb drive t time T absolute temperature tb flexure beam thickness tc out of plane comb finger thickness

Vact electrostatic actuation voltage VCA output amplitude of charge amplifier VDM output amplitide of demodulator VE input amplitude of sinusoidal excitation voltage VLP output of low pass filter VLPMAX maximum output voltage at low pass filter VVA output amplitude of voltage amplifier vy relative motion of proof mass with respect to base Vy relative motion amplitude w flexure beam width Wplate effective plate width ws shin beam width wt thigh beam width x0 initial comb finger gap

x01 small gap distance on top side of the tri-plate comb drive

x02 small gap distance on bottom side of the tri-plate comb drive

ix

x03 larger gap distance on the tri-plate comb drive

XN noise displacement y0 initial comb finger overlap length yB base displacement YB base displacement amplitude ym absolute proof mass displacement α parameter dependant on the WplateLplate ratio ∆C Change in capacitance from a tri-plate drive in a differential

arrangement ε permittivity of free space κ second moment of inertia ratio (shinthigh) ζ damping ratio microair absolute viscosity of air ω base excitation frequency in rads ωn sensor natural frequency in rads ωv electrostatic actuation voltage frequency in rads L1 Sensor fabricated using SoiMUMPS with a lateral comb drive T1 Sensor fabricated using PolyMUMPS with a transverse comb drive T2 Sensor fabricated using SoiMUMPS with a tri-plate comb drive

x

1 INTRODUCTION

11 RESEARCH PURPOSE

The modern automobile consists of many mechanical systems such as power seats

windshield wipers mirrors trunks and windows which are all susceptible to breakdown

Without any condition monitoring system the breakdown is usually catastrophic and

requires an expensive part replacement Real-time condition monitoring allows for early

detection of faults which could require a simple solution such as the application of a

lubricant to fix This prolongs the useful life of the component and prevents sudden and

unexpected failure Real-time condition monitoring can be accomplished by examining the

vibration signature of a mechanical system For example an automobile power window

consists of a DC motor and its associated bearings and couplings a gear reduction system

consisting of worm and spur gears and kinematic links Faults resulting in excessive

vibrations may be caused by coupling misalignment bearing failure or gear train failure

Coupling misalignments occur at the connection between the drive shaft and the driven

shaft and are typically due to imperfect manufacturing Bearing failure is caused by lack

of lubrication or moisture contamination causing rusting while gear train failure is caused

by misaligned broken cracked or chipped gear teeth [1] Each fault occurs at a

characteristic frequency and so the state of the mechanical system can be determined by

monitoring the amplitudes of the relevant frequencies Vibrations due to coupling

misalignments occur at harmonics of the shaft rotational speed Gear vibrations occur at

the gear turn speed or at sidebands of the gear mesh frequency [1] Ball bearing vibrations

1

2

may be caused by outer bearing race defects fOB inner bearing race defects fIB ball

defects fB and train defects fT which all occur at specific frequencies [12]

cos( )12

dBOB r

p

bNf fd

β = minus

cos( )12

dBIB r

p

bNf fd

β = +

2cos( )1

2p r d

Bd p

d f bfb d

β = minus

cos( )12

drT

p

bffd

β = minus

(1)

where NB is the number of balls fr is the rotation frequency bd is the ball diameter β is the

contact angle between the ball and races and dp is the ball pitch diameter By monitoring

the real time vibration signature of the mechanical system anomalies can be quickly

identified and fixed

12 MICROELECTROMECHANICAL SENSORS

Microelectromechanical systems (MEMS) inertial sensors provide a small footprint

with sensitivities that are either comparable or exceed any macro sensor along with the

capability of mass production and low unit cost These sensors utilize compliant micro-

flexures attached to a proof mass that displaces in response to an environmental

acceleration Many transduction mechanisms have been developed that convert the

displacement into a measurable electric signal and include thermal piezoresistive

piezoelectric optical and capacitive methods Most MEMS sensors are silicon based and

are fabricated using surface or bulk micromachining Surface micromachining creates free

standing movable structures on top of a substrate using a combination of sacrificial layers

3

and structural layers which are commonly polysilicon [3] In bulk micromachining the

mechanical structures are defined using a removal process where bulk material typically

silicon is etched away MEMS sensors have been used for vibration and shock monitoring

on industrial systems and robotics guidance and navigation in global positioning systems

(GPS) seismometry in earthquake prediction image stabilization in digital cameras and

automobile safety and stability [4]

13 AUTOMOTIVE SENSORS

Sensors cover every major aspect of a modern automobile power-train sensors

monitor fuel combustion and emissions chassis sensors monitor road traction and tire

condition and body sensors facilitate air-bag deployment and vehicle proximity for radar

guided cruise control [5] Pressure sensors which typically consist of a piezoresistive

strain sensing element attached to a silicon diaphragm that deflects when exposed to an

applied pressure are one of the first micro-machined sensors used in an automobile

Implemented as manifold absolute pressure (MAP) sensors they allow precise control of

the air fuel ratio which allows the catalytic converter to efficiently reduce tailpipe

emissions [6] Variable reluctance sensors based on electromagnetics are used for

automobile traction control and produce a voltage output that is dependant on the

magnetic flux variations between a rotating component and the sensor bias magnet [57]

MEMS based linear accelerometers are utilized for airbag deployment upon impact and

provide a lower cost and smaller package over the older ball-in-tube accelerometers [6]

4

14 OPERATING ENVIRONMENT

Through proper packaging MEMS sensors can withstand the harsh automotive

environment where they face cyclic temperatures (-40degC to 150degC) mechanical shock or

vibration (10g-500g) and exposure to various fluids (oil brake fluid or engine coolant)

and airborne particles [8] Temperature changes induce thermal stresses causing

undesirable bending of the silicon microstructures that changes their dynamic response

Mechanical shock which could occur during a high impact collision or vibration from

uneven road conditions cause the delicate microstructures to fracture at points of stress

concentration These effects could be reduced by efficient design Use of a folded beam

micro-flexure that allows expansion and contraction or use of fillets at corners which

reduce the chance of fracture Bonding pads wire bonding and onboard circuit

components can be protected from air borne particles which could cause shorting using

parylene or gels Alternatively the entire package can be sealed hermetically which also

protects the MEMS device [8] In some cases such as fluid pressure sensors intimate

contact between the MEMS device and potentially corrosive fluid is necessary A recent

approach used a nanometer thin coating of silicon carbide (SiC) over a foundry fabricated

MEMS sensor using low temperature low pressure chemical vapor deposition (LPCVD)

[9] The SiC known for its chemical inertness did not compromise the mechanical

characteristics of the original design Therefore MEMS provides a strong platform for an

automotive sensor

5

15 OBJECTIVES AND CONTRIBUTION

The objective of this work is the development of a low frequency MEMS vibration

sensor The sensor must be low cost and mass producible which can be achieved using

standard foundry fabrication processes such as the Polysilicon Multi-User MEMS

Processes (PolyMUMPS) or the Silicon-on-Insulator Multi-User MEMS Processes

(SoiMUMPS) In addition the number of additional processing steps which are required

if integrated CMOS circuitry is used or if vacuum packaging is required must be

minimized This can be achieved using an off-chip circuit made with widely available

discrete components and sensor operation in air

In addition this work presents a MEMS sensor that is fabricated using standard

foundry processes such as PolyMUMPS and SoiMUMPS compared to many MEMS

sensors in literature that are fabricated using customized processes

16 RESEARCH OUTLINE

This work demonstrates three iterations of a low frequency capacitive vibration

sensor fabricated using the standard foundry processes Each sensor consists of a

suspended proof mass that displaces in response to an external vibration The proof mass

is attached to a folded beam micro-flexure compliant in one direction and its lateral

movement is sensed using a capacitive comb drive The first sensor T1 is fabricated in

PolyMUMPS and uses a transverse capacitive comb drive The second L1 fabricated in

SoiMUMPS uses a fully linear lateral capacitive drive and achieves an increase in

sensitivity from T1 Finally the third sensor T2 is fabricated in SoiMUMPS and uses a

differential tri-plate capacitive drive This sensor takes advantage of the increase in

6

sensitivity gained from a transverse drive but rejects off-axis displacement errors and

common mode using a differential tri-plate arrangement In addition an off-chip

capacitance to voltage readout circuit is fabricated and tested on sensor L1

Chapter 2 begins with an overview on the various methods of fault detection

followed by a review of MEMS vibration sensors each with a different transduction

mechanism Chapter 3 outlines the sensor structure focusing on the mechanical design

and sensing principal Chapter 4 focuses on the readout circuitry layout and simulation

results Chapter 5 outlines the various source of noise including mechanical noise arising

from the sensor and electric noise from the readout circuit Chapter 6 overviews the

foundry fabrication processes used including PolyMUMPs and SoiMUMPS The

experimental results of two sensors T1 and L1 which were fabricated and tested within

the thesis time frame are presented in Chapter 7 Chapter 8 concludes the thesis with a

brief summary and outline for future work

2 LITERATURE REVIEW

21 EXISTING METHODS FOR FAULT DETECTION

As this thesis develops a sensor for fault detection using vibration monitoring this

section overviews three existing methods for fault detection in mechanical systems First

motor current monitoring is discussed as it has the ability to measure significant bearing

faults in motor systems Next temperature sensors specifically thin film thermocouples

(TFTC) and MEMS temperature sensors are discussed These sensors are mounted in the

chamber of a lubricating fluid or in close proximity to a moving component and monitor

temperature fluctuations caused by increased wear Finally acoustic emission (AE)

sensors are discussed which have the ability to detect ultrasonic emissions generated by

plastic deformation in the crystalline lattice of a mechanical component

211 MOTOR CURRENT MONITORING

Monitoring of stator current is used as a method of determining bearing related faults

that produce radial motion between rotor and stator This motion varies the air gap flux

density producing stator currents at predictable frequencies fbng that represent the

bearing faults [10]

bng e vf f m f= plusmn sdot (2)

where fe is the electrical supply frequency m=123hellip and fv is one of the characteristics

frequencies listed in the equations of (1)The bearing damage must be significant to

produce noticeable current fluctuations and it may be difficult to distinguish from current

fluctuations due to voltage imbalances machine saturation or statorarmature faults [2]

7

8

212 TEMPERATURE SENSORS

Temperature sensors are largely used in high temperature high load and high speed

environments typically turbine engines and industrial tools Thermocouple sensors

consist of two dissimilar metals joined together at two junctions and when exposed to a

temperature difference generate a thermoelectric voltage dependant only on the material

properties and junction temperature difference [11] In turbine engines these sensors can

be placed on bearing casings or in the lubricant to monitor temperature fluctuations due

to increased friction The thermocouples are fabricated using bulk materials or micro-

machined thin films (TFTC) with the latter providing a faster response small package

and substantially lower production costs [12] The TFTC can be embedded for in situ

monitoring of processes in hostile environments One study successfully embedded

TFTC into electroplated nickel and attained device characteristics on par with surface

mounted devices [13] More recently MEMS temperature sensors have been developed

that consist of micro-machined semiconductor material that undergoes structural

deformation with temperature changes A common implementation is a bent beam

structure whose temperature induced deflection δ is expressed as [14]

( ) 3

2 2

sin12 cos sinz

EA T LEI AEL

α βδ 2β β

∆=

+

(3)

Temperature sensors are well suited for localized fault monitoring and detection but can

only characterize a mechanical system consisting of multiple components through the use

of multiple sensors

9

213 ACOUSTIC EMISSION SENSORS

Acoustic emission (AE) sensors have been used to characterize wear in machine

tools and monitor bearing and gear problems in centrifugal pumps [1516] First

developed as a Non-Destructive Testing (NDT) technique to detect cracks in civil

structures these sensors detect acoustic emissions generated by the release of vibration

waves in a crystalline lattice due to plastic deformation or crack growth [16]

Measurements are made using piezoelectric transducers with high natural frequencies

100 kHz to 1 MHz to capture the ultrasonic AE emissions An AE sensor is useful as it

has the ability to detect subsurface cracks in gear teeth or bearings before appearing on

the surface causing further damage [17] More recently MEMS acoustic sensors have

been developed and one design includes multiple transducers on a single substrate

which each detect acoustic emission energy at different frequencies [18] This helps

distinguish spurious acoustic emissions arising from impact and friction from those

arising from plastic deformation When compared to typical piezoelectric sensors the

MEMS devices have lower sensitivities and fail to detect some acoustic emissions

[1819] In addition the acoustic emission signal suffers severe attenuation as is crosses

various interfaces such as a gearbox or bearing casings In one experiment consisting of

a pinion gear and an associated bearing a 44dB attenuation was seen between an AE

sensor placed directly on the pinion to one placed on the bearing casing and in some

cases intermediate loss of the signal was observed [20]

10

22 MEMS VIBRATION SENSORS

This thesis focuses on the development of a MEMS vibration sensor for fault

detection as the vibration signature of a mechanical system has the potential to give an

overview of the entire system Many types of MEMS vibration sensors exist with

different operating principals based on resonance the piezoelectric effect and

displacement variation

221 RESONANT SENSORS

Resonant sensors are based on the resonant frequency shift of a fixed-fixed micro-

beam that is excited into resonance using electrostatic thermal or piezoelectric methods

Under an applied axial load there is a shift in the resonant frequency which is expressed

as [21]

2 2

2

1 12 1

nn n

EI Fl2

fl A Eα γ

π ρ= +

I

(4)

where l is the beam length E is Youngrsquos Modulus I is the moment of inertia ρ is the

density A is the cross section area F is the applied axial force αn and γn are mode

dependant coefficients and are 4730 and 0295 for mode 1 in the case discussed The

axial force F is generated by a proof mass attached to one beam end which moves in

response to an external acceleration The performance of resonant sensors are dependant

on the quality factor Q which is the ratio of total energy stored in the system to the

energy lost per cycle and determines the sharpness and amplitude of the resonance peak

A high Q resonator develops a peak that is easily distinguishable in a phase locked loop

control circuit and allows for improved performance and resolution [3] A resonator Q

value is lowered by energy losses to the surrounding fluid attached supports and

11

material Losses to the fluid surrounding the MEMS structure is the dominant loss

mechanism for MEMS sensors operated in air and can be greatly reduced by vacuum

packaging which can be accomplished using glass-silicon anodic bonding [22] In

addition to increasing fabrication cost the long term stability of the vacuum degrades

with ultimate device failure stemming from leakage through micro cracks and defects

and out-gassing [22] Support losses arise from the restoring forces they generate and

can be reduce by balanced designs such as a double ended tuning fork (DETF) or

operation in higher modes [3] In addition the sensitivity of a resonant sensor can be

increased by the addition of micro-levers to increase the axial force In one such design a

two stage micro-lever increased the sensitivity by an order of magnitude over a more

conventional single lever arm [2324]

222 PIEZOELECTRIC SENSORS

Piezoelectric materials produce a charge in response to an applied force This is an

intrinsic effect in materials such as Zinc Oxide (ZnO) whereas materials such as Barium

Titanate or Lead Zirconate Titanate (PZT) need to be poled by placing the material in a

strong electric field at an elevated temperature [3] The piezoelectric effect is expressed

mathematical by [25]

i ij j ik kD d T Eε= + (5)

where D represents the electrical displacement (Cm2) d is the charge coefficient (CN)

T is the applied stress matrix (Nm2) ε is the permittivity matrix (Fm) and E is the

electric field matrix (Vm) The subscripts i=123 j=123456 and k=123 represent

the direction to which the physical property is related For micro-sensors poled in the out

of plane direction only the D3 term is relevant and if fabricated using ZnO or PZT only

12

the charge coefficients d31 d32 and d33 are relevant [25] The piezoelectric material is

commonly deposited as a thin film on the surface of compliant structures in bulk micro-

machined sensors or incorporated as a layer in surface micro-machining on a similar

structure [26] However the inclusion of multiple materials introduces large temperature

variations causing out of plane bending especially in surface micro-machined devices

[2627] Poled piezoelectric materials could be depolarized which serves to reduce the

piezoelectric affect in the material This can occur if the sensor is exposed to a strong

electric field of the opposite polarity high temperatures in excess of the Curie point or

high mechanical stress Also the piezoelectric effect reduces over time an effect that can

be reduced with the addition of composite elements [28] Foundry processes such as

PolyMUMPS and SoiMUMPS do not incorporate piezoelectric layers thus sensors using

this method of transduction must use a customized process The piezoelectric sensor

outlined in [26] uses polysilicon surface micro-machining similar to PolyMUMPS

however it incorporates a ZnO layer that is not available in the foundry process

223 DISPLACEMENT VARIATION SENSORS

Displacement variation sensors consist of a proof mass connected to a compliant

micro-flexure The movement of the proof mass is sensed using optical electron

tunneling or more commonly capacitive methods A recent optical accelerometer

utilized a novel multilayer nano-grating for in-plane displacement sensing of the proof

mass [29] For small space variations between the nano-gratings a large change in the

optical reflection amplitude of the grating was observed [30] Even though optical

sensors offer resolutions approaching the Brownian noise limit they are not available in

small packages that are easily placed in space constricted areas for a low cost [29]

13

The electron tunneling effect is observed when the proximity between two electrodes

one flexible and the other fixed is on the order of 10Ǻ [31] If the two electrodes are

biased with voltage VB a tunneling current It flows between the gap [31]

exp( )t B I tgI V xαinfin sdot minus Φ (6)

where α1 is a constant and equal to 1025 Ǻ-1 eV-frac12 Ф is the effective height of the

tunneling barrier and xtg is the tunneling gap When the flexible electrode moves in

response to acceleration a feedback controller maintains the original distance while

determining the applied acceleration Electron tunneling sensors are able to sense

accelerations in the microg-ng range due to the exponential relation between tunneling current

and gap However these highly sensitive devices are susceptible to thermo-mechanical

noise which could be reduced by operation at low pressure [32] It is also important for

the two electrodes to be coated with metal which is difficult to achieve with foundry

fabrication

Capacitive sensors are based on the parallel plate capacitor and are implemented

using inter-digitated comb fingers in an in-plane lateral or transverse drive or in an out of

plane parallel plate drive [33] These sensors have the advantage of high sensitivity low

power consumption small package low temperature dependence and are easily

integrated with existing foundry fabrication processes [34] The lateral drive offers a fully

linear response but low sensitivity whereas the transverse and parallel plate drives

achieve high sensitivity if gap distances are of a few microns and large capacitance areas

are utilized However both drives have a non-linear response and are prone to pull-in

Also if these sensors are fabricated using a surface micromachining process such as

PolyMUMPS where the ground plane is a few microns above the structural layer they

14

are prone to electrical failure due to shorting This was observed in an experiment where

similar sensors were placed in a vibrating environment with a peak of 120g and shorting

between the structural layer and ground plane was found to be the predominate mode of

failure [35] However with certain fabrication processes such as SoiMUMPS this issue

can be neglected as the substrate material below any moving component is removed

Most capacitance sensors are fabricated using an integrated CMOS (Complementary

Metal Oxide Semiconductor)-MEMS technique where the MEMS sensor is fabricated

before during or after the CMOS circuit fabrication [36] Although these sensors offer

high-sensitivity and small parasitic capacitance they suffer from in-plane and out of

plane curling of the beam which reduces the capacitance between adjacent comb fingers

[37] In addition multiple processing steps are required for device fabrication

More than the other transduction methods MEMS capacitive displacement sensors are

attractive for practical implementation in an automobile and therefore is the selected

method for the vibration sensor presented in this work Their low temperature dependence

is an ideal characteristic for an automobile environment that faces varying temperature If

implemented with off-chip readout circuitry they require a standard foundry fabrication

process which provides a strong platform for mass production and low unit cost

3 SENSOR STRUCTURE

This chapter presents the physical structure of the three MEMS vibration sensors

which includes the displacement sensing capacitive comb drive and the micro-

flexureproof mass system First the capacitive comb drive and the three variations

implemented the transverse drive lateral drive and differential tri-plate drive are

discussed Next the mechanical structure is presented and includes the micro-flexure

selection quality factor and equations of motion Finally instability that places a limit

on the operating range of the sensor as a result of a gap closing capacitive comb drive is

discussed

31 SENSING

Capacitive sensors are based on the parallel plate capacitor and are commonly

implemented using inter-digitated comb fingers A capacitance can be realized if a voltage

is placed between two closely spaced plates If the fringing electric field is neglected the

capacitance of the parallel plate capacitor Cpp (F) is expressed as [28]

air pppp

k AC

dε

= (7)

where kair=1 is the dielectric constant of air at atmospheric pressure ε=88542times10-12 F m-1

is the permittivity of free space App is the overlapping area (m2) and d is the distance

between the two plates (m)

15

16

Figure 3-1 An inter-digitated comb drive where changes in capacitance are generated by either changes in gap distance x0 or in the overlap area y0timest

Figure 3-1

For a set of inter-digitated comb fingers or comb drive shown in the total

capacitance Ctotal is expressed as [28]

0 0

0 0

( ) ( )( ) (Total cy y y yC N t

)x x x xε

+ += + + minus

(8)

where N is the number of comb fingers on the movable comb y0 is the initial comb finger

overlap length (m) tc is the out of plane comb finger thickness (m) and x0 is the inital

finger gap (m) In this work the comb drive is implemented in three arrangements

transverse lateral and differential tri-plate In a transverse comb drive changes in

capacitance are generated by transverse movements along the x-axis resulting in changes

to the gap x0 between the movable and fixed comb fingers The nonlinear relation

between total capacitance CTotalT and displacement is expressed as

17

00 0

1 1( ) (Total TC N ty

)x x x xε

= + minus +

(9)

Although the drive sensitivity part is not constant an average sensitivity can be

assumed for small displacements [38]

Total TC xpart

In a lateral comb drive changes in capacitance are generated by lateral movements

along the y-axis resulting in changes to the overlapping area y0timestc The total capacitance

CTotalL between all comb fingers is expressed as

00

2 ( )Total LN tC yx

yε= +

(10)

The sensitivity of the lateral comb drive SL is expressed as

0

2Total LL

C N tSy x

εpart= =

part

(11)

The lateral comb drive was first introduced by W C Tang T H Nguyen and R T

Howe [33] in an electrostatic resonator and has since been used numerous times in both

drive and sense configurations It has the advantage of a fully linear response and

negligible instability but offers low sensitivity The sensitivity for the lateral drive L1

implemented in this work is 0038 pF microm-1

A differential tri-plate drive shown in Figure 3-2 which is a variation of the

transverse drive is implemented to improve sensitivity and reduce noise [39]

18

Figure 3-2 An inter-digitated comb drive in a differential tri-plate arrangement where x01 x02 and x03 are the gap distances and x01= x02ltlt x03 to improve sensitivity

In the tri-plate drive the at rest gap distances between the two capacitors formed by any

movable finger is not equivalent For a differential arrangement insuring that x01=

x02ltltx03 the linearity for small displacements (less than 1 microm) is enhanced over the case

of a single ended arrangement where x01=x02=x03=x0 which is equivalent to the

transverse drive This is shown graphically in Figure 3-3 where x01= x02=x03=2 microm in

the lower curve and x01= x02= 2 microm and x03= 15 microm in the upper curve A substantial

improvement in linearity is attained with the tri-plate drive

19

0

005

01

015

02

025

03

035

0 02 04 06 08 1

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-3 Comparison of the effect of gap distances on linearity for two case lsquorsquo represents case 1 where the fingers in are implemented in a differential tri-plate arrangement and x01= x02= 2 microm and x03= 15 microm lsquorsquo represents the case 2 where the fingers are implemented in a single ended arrangement and x01= x02=x03=2 microm

The tri-plate drive is also implemented in a differential arrangement where the output

is the change in capacitance rather than the total capacitance resulting from an external

excitation This reduces errors cause by off-axis excitation and common mode noise The

output of the differential tri-plate drive is expressed as [39]

Total Top Total BottomC C C∆ = minus where

003 01

1 1( ) (Total TopC N ty

)x x x xε

= +

minus +

(12)

20

002 03

1 1( ) (Total BottomC N ty

)x x x xε

= +

minus +

A graphic comparison of the capacitancedisplacement relation for sensors T1 and T2 is

shown in Figure 3-4 For these gap closing drives it is assumed that for displacements

less that 1 microm the relation is linear indicated by the dashed line above each curve The

sensitivity within this region is approximately 0046 pF microm-1 and 0253 pF microm-1 for T1

and T2 respectively The order of magnitude improvement is attributed to the differential

tri-plate arrangement and thicker comb finger

0

02

04

06

08

1

12

14

16

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-4 lsquorsquo represents the output of the differential tri-plate drive while lsquorsquo represents the output of the transverse drive The dashed line represents the linear approximation for displacements less that 1 microm For the purpose of comparison the output of the transverse drive was equalized to zero at zero displacement

21

In this work the transverse drive is first implemented in sensor T1 using the

PolyMUMPS fabrication process Then in an effort to linearize the sensor response the

lateral drive is used in L1 using SoiMUMPS The reduction in drive sensitivity due to the

lateral drive is further reduced by increasing the gap distance to 3 microm which is

recommended by the foundry to ensure a successful device These reductions are

balanced by increasing the number of comb fingers N and selecting a substrate thickness

that is 125 times large than T1 The third device T2 is fabricated using SoiMUMPS and

uses the tri-plate differential drive This design capitalizes on the increase in sensitivity of

a gap closing drive and larger substrate thickness attaining an order of magnitude

increase in sensitivity over T1 and L1 The specifications for each drive along with the

successive improvement in drive sensitivity are outlined in Table 3-1

Table 3-1 Specifications for each capacitive comb drive implemented in this work T1 L1 and T2

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Comb Finger Thickness tc (microm) 2 25 25 Gap Distance x0 (microm) 2 3 2 Number of Fingers N 200 256 30 Overlap length y0 (microm) 45 10 70 Drive Sensitivity (pF microm-1) 0046 0038 0273

32 MECHANICAL ANALYSIS

321 MICRO-FLEXURE SELECTION

The sensor structure consists of a proof mass connected to a compliant micro-flexure

that itself is anchored to the sensor substrate The proof mass displaces in response to an

22

applied acceleration and this displacement is sensed using the capacitive comb drive

attached to the proof mass

Proof mass

Folded beam micro-flexure

Anchor

Figure 3-5 Sensor Structure

This section focuses on micro-flexure selection as it should exhibit a large linear

displacement range and have a high lateral stiffness ratio to reduce cross-axis sensitivity

A number of micro-flexures have been used in MEMS research these include the

clamped-clamped crab-leg and folded beam shown in Figure 3-6-through 3-7

The clamped-clamped flexure exhibits a high stiffness ratio however for large lateral

displacements extensional axial forces buildup within the beams resulting in a non-linear

force-displacement relation [40] Within the linear regime the stiffness of the clamped-

clamped flexure is expressed as [40]

3

3b

yEt wk

L=

(13)

23

where E is the Youngrsquos Modulus of the beam material (N m-2) tb is the beam out of

plane thickness (m) w is the beam width (m) and L is the beam length (m) The stiffness

ratio is expressed as [40]

2x

y

k Lk w

=

(14)

If L is two orders of magnitude greater than w this results in a ratio of approximately

10000 The linear displacement range is determined using Finite Element Analysis (FEA)

of a clamped-clamped flexure with L=387 microm w=4 microm and tb=2 microm The result of this

simulation is plotted in and compared to a structure with constant stiffness ky

= 035 N m-1 predicted by Eq (13) The linear relation is valid for small deflections that

are less than 2 microm In addition this micro-flexure is susceptible to buckling as a result of

a compressive residual stress that may arise due to the fabrication process [41]

Figure 3-9

In the crab leg design the thigh section is meant to relieve the extensional axial

stress postponing the occurrence of non-linearity until greater lateral displacement [42]

The stiffness of the crab leg flexure is expressed as [43]

3

44

b s s ty

s s t

Et w L LkL L L

κκ

+= +

(15)

where κ=IsIt=(wswt)3 is a ratio of the second moment of inertia of the shin (s) and thigh

(t) ws and wt represent the width and Ls and Lt represent the length of the shin and thigh

regions respectively The stiffness ratio is expressed as [43]

34

4t s s tx

y s t s

w L L Lkk w L L L

αα t

+= +

(16)

The linear displacement range of a crab leg flexure with Ls=285 microm Lt=30 microm ws=wt=3

microm and tb=2 microm is determined using FEA The results are plotted in Figure 3-9 and

compared to a structure of constant stiffness ky=035 N m-1 predicted by Eq (15) The

24

non-linearity is greatly reduced when compared to the clamp-clamped flexure with Eq

accurately predicting the deflection until 4 microm However the flexure is more

susceptible to off-axis deflection as there is an order of magnitude reduction in the

stiffness ratio as compared to the clamped-clamped flexure

(15)

The folded beam flexure stands out as it is designed to exhibit a large linear

displacement range and to relieve residual stress due to the manufacturing process by

allowing the beams to expand and contract along the axial direction [40] It was first used

in an electrostatic resonator and has subsequently been used for electrostatic actuation in

a magnetic hard drives and micro-positioning [3344] If the truss section is rigid

ItrussgtgtIbeam then the stiffness of the flexure is expressed as [4043]

3

32b

yEt wk

L=

(17)

The stiffness ratio is expressed as [40]

2x

y

k Lk w

=

(18)

Figure 3-6 Clamped-clamped flexure

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

LIST OF FIGURES

FIGURE 3-1 AN INTER-DIGITATED COMB DRIVE WHERE CHANGES IN CAPACITANCE ARE GENERATED BY

EITHER CHANGES IN GAP DISTANCE X0 OR IN THE OVERLAP AREA Y0timesT 16 FIGURE 3-2 AN INTER-DIGITATED COMB DRIVE IN A DIFFERENTIAL TRI-PLATE ARRANGEMENT WHERE X01

X02 AND X03 ARE THE GAP DISTANCES AND X01= X02ltlt X03 TO IMPROVE SENSITIVITY 18 FIGURE 3-3 COMPARISON OF THE EFFECT OF GAP DISTANCES ON LINEARITY FOR TWO CASE lsquorsquo REPRESENTS

CASE 1 WHERE THE FINGERS IN ARE IMPLEMENTED IN A DIFFERENTIAL TRI-PLATE ARRANGEMENT AND X01= X02= 2 ΜM AND X03= 15 ΜM lsquorsquo REPRESENTS THE CASE 2 WHERE THE FINGERS ARE IMPLEMENTED IN A SINGLE ENDED ARRANGEMENT AND X01= X02=X03=2 ΜM 19

FIGURE 3-4 lsquorsquo REPRESENTS THE OUTPUT OF THE DIFFERENTIAL TRI-PLATE DRIVE WHILE lsquorsquo REPRESENTS THE OUTPUT OF THE TRANSVERSE DRIVE THE DASHED LINE REPRESENTS THE LINEAR APPROXIMATION FOR DISPLACEMENTS LESS THAT 1 ΜM FOR THE PURPOSE OF COMPARISON THE OUTPUT OF THE TRANSVERSE DRIVE WAS EQUALIZED TO ZERO AT ZERO DISPLACEMENT 20

FIGURE 3-5 SENSOR STRUCTURE 22 FIGURE 3-6 CLAMPED-CLAMPED FLEXURE 24 FIGURE 3-7 CRAB LEG FLEXURE 25 FIGURE 3-8 FOLDED BEAM FLEXURE 25 FIGURE 3-9 COMPARISON OF THE NON-LINEAR DEFLECTION AMONG THE THREE MICRO-FLEXURES

CONSIDERED CLAMPED-CLAMPED CRAB-LEG AND FOLDED BEAM ALL FLEXURES WERE DESIGNED TO HAVE THE SAME STIFFNESS OF 035 N M-1 26

FIGURE 3-10 LUMPED MASS APPROXIMATION WITH BASE EXCITATION YB(T) 27 FIGURE 3-11 LUMPED MASS APPROXIMATION WITH DIRECT MASS EXCITATION Z(T) 28 FIGURE 3-12 GRAPH REPRESENTING STABILITY REGIONS FOR GAP CLOSING CAPACITIVE DRIVES lsquorsquo

REPRESENTS THE RHS OF EQ (32) FOR T2 WHILE lsquorsquo REPRESENTS THE RHS OF T1 34 FIGURE 4-1 IDEAL CHARGE AMPLIFIER 36 FIGURE 4-2 CAPACITANCE TO VOLTAGE READOUT CIRCUIT WHICH IS MADE UP OF FOUR ELEMENTS THE

CHARGE AMPLIFIER VOLTAGE AMPLIFIER DEMODULATOR AND LOW PASS FILTER 37 FIGURE 4-3 ORCAD SIMULATION RESULTS OF THE CAPACITANCE TO VOLTAGE READOUT CIRCUIT THE

EXCITATION VOLTAGE VE IS 75 VPK AND THE OUTPUT OF THE VOLTAGE AMPLIFIER VVA IS APPROXIMATELY 13125 VPK 38

FIGURE 4-4 ORCAD SIMULATION RESULTS SHOWING THE SINUSOIDAL OUTPUT OF THE DEMODULATOR VDM WITH A 984 VPK-PK AMPLITUDE AND THE DC OUTPUT OF THE LOW PASS FILTER VLP AT APPROXIMATELY 5 V 39

FIGURE 5-1 ORCAD SIMULATION OF THE NOISE VOLTAGE SPECTRAL DENSITY ENO AT THE OUTPUT OF THE VOLTAGE AMPLIFIER 42

FIGURE 5-2 TOTAL OUTPUT NOISE ENO AT THE VOLTAGE AMPLIFIER WHICH IS EXPRESSED AS ENO=( int ENO2)12

42 FIGURE 6-1 POLYMUMPS FABRICATION PROCESS FOR A SIMPLE CANTILEVER BEAM 44 FIGURE 6-2 SOIMUMPS FABRICATION PROCESS FOR A SIMPLE CANTILEVER BEAM 45 FIGURE 7-1 THE SOLID LINE REPRESENTS THE THEORETICAL CAPACITANCE VERSUS DISPLACEMENT

RELATION WHILE THE SOLID POINTS REPRESENT THE EXPERIMENTAL VALUES THE DASHED LINE REPRESENTS THE ASSUMED LINEAR RELATION FOR DISPLACEMENTS LESS THAN 1ΜM THE EXPERIMENTAL AVERAGE DRIVE SENSITIVITY IS APPROXIMATELY 0024 PF ΜM-1 FOR DISPLACEMENTS GREATER THAN 12 ΜM DRIVE INSTABILITY IS OBSERVED AND IS ATTRIBUTED TO THE PULL-IN EFFECT 47

FIGURE 7-2 lsquotimesrsquo REPRESENTS THE AMPLITUDE AND SOLID LINE REPRESENTS THE PHASE RESPONSE THE RESONANCE PEAK IS LOCATED AT 3600 HZ AND ACCOMPANIED BY A PHASE SHIFT OF 168deg 48

FIGURE 7-3 A SEM MICRO-GRAPH SHOWING THE SENSOR TI FABRICATED USING POLYMUMPS THE FOUR FOLDED BEAM FLEXURES HAD A LENGTH OF 650 ΜM AND TOTAL STIFFNESS OF 14 NM THE EFFECTIVE PROOF MASS WAS 29E-9 KG AND THE SYSTEM HAD A NATURAL FREQUENCY OF 3500 HZ 50

FIGURE 7-4 A SEM MICRO-GRAPH OF THE TRANSVERSE COMB DRIVE IMPLEMENTED IN T1 WITH 200 COMB FINGERS A GAP DISTANCE OF 2 ΜM OVERLAP LENGTH OF 45 ΜM AND THICKNESS OF 2 ΜM THE

vi

AVERAGE EXPERIMENTAL SENSITIVITY WAS FOUND TO BE 0024 PF ΜM-1 FOR DISPLACEMENTS OF LESS THAN 1 ΜM 50

FIGURE 7-5 THE SOLID LINE REPRESENTS THE THEORETICAL CAPACITANCE TO VOLTAGE RELATION EXPRESSED BY EQ (37) WITH A SENSITIVITY OF 141 V PF-1 THE SOLID POINTS REPRESENT THE EXPERIMENTAL DATA POINTS WHEN DISCRETE CAPACITORS 1-4 PF ARE USED IN THE READOUT CIRCUIT THE HORIZONTAL DASHED LINE REPRESENTS THE OUTPUT VOLTAGE OF THE MEMS SENSOR AT REST AND CORRESPONDS TO A CAPACITANCE OF 35 PF 52

FIGURE 7-6 OSCILLOSCOPE TRACE OF THE INPUT EXCITATION VOLTAGE VE asymp 75 VPK AND OUTPUT OF THE VOLTAGE AMPLIFIER VVA asymp 103 VPK READINGS WERE TAKEN WITH THE MEMS SENSOR AT REST 53

FIGURE 7-7 OSCILLOSCOPE TRACE OF THE DEMODULATOR OUTPUT VDM asymp 76 VPK-PK AND OUTPUT OF THE LOW PASS FILTER VLP asymp 36 V READINGS WERE TAKEN WITH THE MEMS SENSOR AT REST 53

FIGURE 7-8 lsquorsquo REPRESENTS THE EXPERIMENTAL DATA POINTS WITH THE DASHED LINE REPRESENTING THE FITTED CURVE SOLID LINE REPRESENTS THE THEORETICAL DATA POINTS ACCORDING TO EQ (10) 54

FIGURE 7-9 EXPERIMENT SETUP WHICH INCLUDES THE MEMS SENOR AND READOUT BOARD MOUNTED ON THE CLAMPED-FREE BEAM AND TEST EQUIPMENT 55

FIGURE 7-10 CLOSE-UP OF THE MEMS SENSOR AND READOUT BOARD MOUNTED ON THE TEST BEAM ALONG WITH THE COMMERCIAL PIEZOELECTRIC ACCELEROMETER ONLY THE CHARGE AMPLIFIER (CA) AND THE VOLTAGE AMPLIFIER (VA) ARE MOUNTED ON THE SAME PROTOTYPE BOARD AS THE MEMS SENSOR DUE TO SPACE CONSTRAINTS 55

FIGURE 7-11 FIGURE COMPARING THE RESPONSE OF A COMMERCIAL PIEZOELECTRIC ACCELEROMETER AND MEMS SENSOR (L1) TO A BEAM VIBRATION A) DYNAMIC RESPONSE OF THE PIEZOELECTRIC ACCELEROMETER REPRESENTING THE TRUE BEAM VIBRATION WITH A PEAK VOLTAGE 114 V CORRESPONDING TO AN APPLIED ACCELERATION OF 114 G B) DYNAMIC RESPONSE OF THE MEMS SENSOR WITH A PEAK VOLTAGE OF 006 V 56

FIGURE 7-12 SENSOR DYNAMIC RESPONSE TO THE VIBRATION OF THE CLAMPED-FREE BEAM AND REPRESENTS A DIRECT EXPORT FROM THE DYNAMIC SIGNAL ANALYZER (AGILENT 35670) TOP TRACE IS THE TIME DOMAIN SIGNAL OF THE LOW PASS FILTER OUTPUT WHERE VLPMAXasymp73MV WHICH REPRESENTS A MAXIMUM ACCELERATION OF APPROXIMATELY 15G THE BOTTOM TRACE REPRESENTS THE FREQUENCY SPECTRUM WITH A PEAK AT 23HZ WHICH CORRESPONDS TO THE BEAM NATURAL FREQUENCY 57

FIGURE 7-13 EXPERIMENTAL RESULTS SHOWING THE RESPONSE OF THE SENSOR TO VARIOUS ACCELERATIONS IN THE RANGE OF 0-20 G WHERE lsquorsquo REPRESENTS THE EXPERIMENTAL DATA POINTS AND THE SOLID LINE REPRESENTS THE FITTED CURVE THE EXPERIMENTAL SENSITIVITY WAS FOUND TO BE 4258 MV G-1 58

vii

NOMENCLATURE a applied acceleration A applied acceleration amplitude App parallel plate overlapping area Asldie-film effective area for slide-film damping c coefficient of damping Cf feedback capacitance Cpp capacitance of the parallel plate capacitor cslide-film coefficient of damping as a result of slide-film damping

csqueeze-film coefficient of damping as a result of squeeze-film damping CTotal total capacitance of a comb drive CTotalBottom Total capacitance of the bottom drive in a tri-plate comb drive CTotalL total capacitance of a lateral drive CTotalT total capacitance of a transverse drive CTotalTop Total capacitance of the top drive in a tri-plate comb drive d distance between the parallel plates E Youngrsquos Modulus of sensor material eno noise voltage spectral density Eno total output noise voltage EnoAMP total output amplifier noise voltage EnoR total output resistor noise voltage

f excitation frequency in Hz fn system natural frequency in Hz

Fact electrostatic actuation force

FN noise force Fx-Elect total electrostatic force for pull-in instability Itruss second moment of inertia of truss section on the folded beam flexure Ibeam second moment of inertia of beam section on the folded beam

flexure kair dielectric constant of air at atmospheric pressure kB Boltzmannrsquos constant kT total stiffness of parallel folded beam flexure kx stiffness of various flexures in x-direction ky stiffness of various flexures in y-direction L flexure beam length Lplate effective plate length

viii

Ls shin beam length Lt thigh beam length m effective proof mass N number of comb fingers on the movable comb

NB number of balls

Qtotal total quality factor

Qf energy lost to surrounding fluid

Qs energy lost through supports

Qm energy lost through sensor material S theoretical sensitivity of sensor SL sensitivity of a lateral comb drive SM theoretical mechanical sensitivity of the proof massflexure system SRO sensitivity of readout circuit ST sensitivity of a transverse comb drive STAVG average sensitivity of a transverse comb drive t time T absolute temperature tb flexure beam thickness tc out of plane comb finger thickness

Vact electrostatic actuation voltage VCA output amplitude of charge amplifier VDM output amplitide of demodulator VE input amplitude of sinusoidal excitation voltage VLP output of low pass filter VLPMAX maximum output voltage at low pass filter VVA output amplitude of voltage amplifier vy relative motion of proof mass with respect to base Vy relative motion amplitude w flexure beam width Wplate effective plate width ws shin beam width wt thigh beam width x0 initial comb finger gap

x01 small gap distance on top side of the tri-plate comb drive

x02 small gap distance on bottom side of the tri-plate comb drive

ix

x03 larger gap distance on the tri-plate comb drive

XN noise displacement y0 initial comb finger overlap length yB base displacement YB base displacement amplitude ym absolute proof mass displacement α parameter dependant on the WplateLplate ratio ∆C Change in capacitance from a tri-plate drive in a differential

arrangement ε permittivity of free space κ second moment of inertia ratio (shinthigh) ζ damping ratio microair absolute viscosity of air ω base excitation frequency in rads ωn sensor natural frequency in rads ωv electrostatic actuation voltage frequency in rads L1 Sensor fabricated using SoiMUMPS with a lateral comb drive T1 Sensor fabricated using PolyMUMPS with a transverse comb drive T2 Sensor fabricated using SoiMUMPS with a tri-plate comb drive

x

1 INTRODUCTION

11 RESEARCH PURPOSE

The modern automobile consists of many mechanical systems such as power seats

windshield wipers mirrors trunks and windows which are all susceptible to breakdown

Without any condition monitoring system the breakdown is usually catastrophic and

requires an expensive part replacement Real-time condition monitoring allows for early

detection of faults which could require a simple solution such as the application of a

lubricant to fix This prolongs the useful life of the component and prevents sudden and

unexpected failure Real-time condition monitoring can be accomplished by examining the

vibration signature of a mechanical system For example an automobile power window

consists of a DC motor and its associated bearings and couplings a gear reduction system

consisting of worm and spur gears and kinematic links Faults resulting in excessive

vibrations may be caused by coupling misalignment bearing failure or gear train failure

Coupling misalignments occur at the connection between the drive shaft and the driven

shaft and are typically due to imperfect manufacturing Bearing failure is caused by lack

of lubrication or moisture contamination causing rusting while gear train failure is caused

by misaligned broken cracked or chipped gear teeth [1] Each fault occurs at a

characteristic frequency and so the state of the mechanical system can be determined by

monitoring the amplitudes of the relevant frequencies Vibrations due to coupling

misalignments occur at harmonics of the shaft rotational speed Gear vibrations occur at

the gear turn speed or at sidebands of the gear mesh frequency [1] Ball bearing vibrations

1

2

may be caused by outer bearing race defects fOB inner bearing race defects fIB ball

defects fB and train defects fT which all occur at specific frequencies [12]

cos( )12

dBOB r

p

bNf fd

β = minus

cos( )12

dBIB r

p

bNf fd

β = +

2cos( )1

2p r d

Bd p

d f bfb d

β = minus

cos( )12

drT

p

bffd

β = minus

(1)

where NB is the number of balls fr is the rotation frequency bd is the ball diameter β is the

contact angle between the ball and races and dp is the ball pitch diameter By monitoring

the real time vibration signature of the mechanical system anomalies can be quickly

identified and fixed

12 MICROELECTROMECHANICAL SENSORS

Microelectromechanical systems (MEMS) inertial sensors provide a small footprint

with sensitivities that are either comparable or exceed any macro sensor along with the

capability of mass production and low unit cost These sensors utilize compliant micro-

flexures attached to a proof mass that displaces in response to an environmental

acceleration Many transduction mechanisms have been developed that convert the

displacement into a measurable electric signal and include thermal piezoresistive

piezoelectric optical and capacitive methods Most MEMS sensors are silicon based and

are fabricated using surface or bulk micromachining Surface micromachining creates free

standing movable structures on top of a substrate using a combination of sacrificial layers

3

and structural layers which are commonly polysilicon [3] In bulk micromachining the

mechanical structures are defined using a removal process where bulk material typically

silicon is etched away MEMS sensors have been used for vibration and shock monitoring

on industrial systems and robotics guidance and navigation in global positioning systems

(GPS) seismometry in earthquake prediction image stabilization in digital cameras and

automobile safety and stability [4]

13 AUTOMOTIVE SENSORS

Sensors cover every major aspect of a modern automobile power-train sensors

monitor fuel combustion and emissions chassis sensors monitor road traction and tire

condition and body sensors facilitate air-bag deployment and vehicle proximity for radar

guided cruise control [5] Pressure sensors which typically consist of a piezoresistive

strain sensing element attached to a silicon diaphragm that deflects when exposed to an

applied pressure are one of the first micro-machined sensors used in an automobile

Implemented as manifold absolute pressure (MAP) sensors they allow precise control of

the air fuel ratio which allows the catalytic converter to efficiently reduce tailpipe

emissions [6] Variable reluctance sensors based on electromagnetics are used for

automobile traction control and produce a voltage output that is dependant on the

magnetic flux variations between a rotating component and the sensor bias magnet [57]

MEMS based linear accelerometers are utilized for airbag deployment upon impact and

provide a lower cost and smaller package over the older ball-in-tube accelerometers [6]

4

14 OPERATING ENVIRONMENT

Through proper packaging MEMS sensors can withstand the harsh automotive

environment where they face cyclic temperatures (-40degC to 150degC) mechanical shock or

vibration (10g-500g) and exposure to various fluids (oil brake fluid or engine coolant)

and airborne particles [8] Temperature changes induce thermal stresses causing

undesirable bending of the silicon microstructures that changes their dynamic response

Mechanical shock which could occur during a high impact collision or vibration from

uneven road conditions cause the delicate microstructures to fracture at points of stress

concentration These effects could be reduced by efficient design Use of a folded beam

micro-flexure that allows expansion and contraction or use of fillets at corners which

reduce the chance of fracture Bonding pads wire bonding and onboard circuit

components can be protected from air borne particles which could cause shorting using

parylene or gels Alternatively the entire package can be sealed hermetically which also

protects the MEMS device [8] In some cases such as fluid pressure sensors intimate

contact between the MEMS device and potentially corrosive fluid is necessary A recent

approach used a nanometer thin coating of silicon carbide (SiC) over a foundry fabricated

MEMS sensor using low temperature low pressure chemical vapor deposition (LPCVD)

[9] The SiC known for its chemical inertness did not compromise the mechanical

characteristics of the original design Therefore MEMS provides a strong platform for an

automotive sensor

5

15 OBJECTIVES AND CONTRIBUTION

The objective of this work is the development of a low frequency MEMS vibration

sensor The sensor must be low cost and mass producible which can be achieved using

standard foundry fabrication processes such as the Polysilicon Multi-User MEMS

Processes (PolyMUMPS) or the Silicon-on-Insulator Multi-User MEMS Processes

(SoiMUMPS) In addition the number of additional processing steps which are required

if integrated CMOS circuitry is used or if vacuum packaging is required must be

minimized This can be achieved using an off-chip circuit made with widely available

discrete components and sensor operation in air

In addition this work presents a MEMS sensor that is fabricated using standard

foundry processes such as PolyMUMPS and SoiMUMPS compared to many MEMS

sensors in literature that are fabricated using customized processes

16 RESEARCH OUTLINE

This work demonstrates three iterations of a low frequency capacitive vibration

sensor fabricated using the standard foundry processes Each sensor consists of a

suspended proof mass that displaces in response to an external vibration The proof mass

is attached to a folded beam micro-flexure compliant in one direction and its lateral

movement is sensed using a capacitive comb drive The first sensor T1 is fabricated in

PolyMUMPS and uses a transverse capacitive comb drive The second L1 fabricated in

SoiMUMPS uses a fully linear lateral capacitive drive and achieves an increase in

sensitivity from T1 Finally the third sensor T2 is fabricated in SoiMUMPS and uses a

differential tri-plate capacitive drive This sensor takes advantage of the increase in

6

sensitivity gained from a transverse drive but rejects off-axis displacement errors and

common mode using a differential tri-plate arrangement In addition an off-chip

capacitance to voltage readout circuit is fabricated and tested on sensor L1

Chapter 2 begins with an overview on the various methods of fault detection

followed by a review of MEMS vibration sensors each with a different transduction

mechanism Chapter 3 outlines the sensor structure focusing on the mechanical design

and sensing principal Chapter 4 focuses on the readout circuitry layout and simulation

results Chapter 5 outlines the various source of noise including mechanical noise arising

from the sensor and electric noise from the readout circuit Chapter 6 overviews the

foundry fabrication processes used including PolyMUMPs and SoiMUMPS The

experimental results of two sensors T1 and L1 which were fabricated and tested within

the thesis time frame are presented in Chapter 7 Chapter 8 concludes the thesis with a

brief summary and outline for future work

2 LITERATURE REVIEW

21 EXISTING METHODS FOR FAULT DETECTION

As this thesis develops a sensor for fault detection using vibration monitoring this

section overviews three existing methods for fault detection in mechanical systems First

motor current monitoring is discussed as it has the ability to measure significant bearing

faults in motor systems Next temperature sensors specifically thin film thermocouples

(TFTC) and MEMS temperature sensors are discussed These sensors are mounted in the

chamber of a lubricating fluid or in close proximity to a moving component and monitor

temperature fluctuations caused by increased wear Finally acoustic emission (AE)

sensors are discussed which have the ability to detect ultrasonic emissions generated by

plastic deformation in the crystalline lattice of a mechanical component

211 MOTOR CURRENT MONITORING

Monitoring of stator current is used as a method of determining bearing related faults

that produce radial motion between rotor and stator This motion varies the air gap flux

density producing stator currents at predictable frequencies fbng that represent the

bearing faults [10]

bng e vf f m f= plusmn sdot (2)

where fe is the electrical supply frequency m=123hellip and fv is one of the characteristics

frequencies listed in the equations of (1)The bearing damage must be significant to

produce noticeable current fluctuations and it may be difficult to distinguish from current

fluctuations due to voltage imbalances machine saturation or statorarmature faults [2]

7

8

212 TEMPERATURE SENSORS

Temperature sensors are largely used in high temperature high load and high speed

environments typically turbine engines and industrial tools Thermocouple sensors

consist of two dissimilar metals joined together at two junctions and when exposed to a

temperature difference generate a thermoelectric voltage dependant only on the material

properties and junction temperature difference [11] In turbine engines these sensors can

be placed on bearing casings or in the lubricant to monitor temperature fluctuations due

to increased friction The thermocouples are fabricated using bulk materials or micro-

machined thin films (TFTC) with the latter providing a faster response small package

and substantially lower production costs [12] The TFTC can be embedded for in situ

monitoring of processes in hostile environments One study successfully embedded

TFTC into electroplated nickel and attained device characteristics on par with surface

mounted devices [13] More recently MEMS temperature sensors have been developed

that consist of micro-machined semiconductor material that undergoes structural

deformation with temperature changes A common implementation is a bent beam

structure whose temperature induced deflection δ is expressed as [14]

( ) 3

2 2

sin12 cos sinz

EA T LEI AEL

α βδ 2β β

∆=

+

(3)

Temperature sensors are well suited for localized fault monitoring and detection but can

only characterize a mechanical system consisting of multiple components through the use

of multiple sensors

9

213 ACOUSTIC EMISSION SENSORS

Acoustic emission (AE) sensors have been used to characterize wear in machine

tools and monitor bearing and gear problems in centrifugal pumps [1516] First

developed as a Non-Destructive Testing (NDT) technique to detect cracks in civil

structures these sensors detect acoustic emissions generated by the release of vibration

waves in a crystalline lattice due to plastic deformation or crack growth [16]

Measurements are made using piezoelectric transducers with high natural frequencies

100 kHz to 1 MHz to capture the ultrasonic AE emissions An AE sensor is useful as it

has the ability to detect subsurface cracks in gear teeth or bearings before appearing on

the surface causing further damage [17] More recently MEMS acoustic sensors have

been developed and one design includes multiple transducers on a single substrate

which each detect acoustic emission energy at different frequencies [18] This helps

distinguish spurious acoustic emissions arising from impact and friction from those

arising from plastic deformation When compared to typical piezoelectric sensors the

MEMS devices have lower sensitivities and fail to detect some acoustic emissions

[1819] In addition the acoustic emission signal suffers severe attenuation as is crosses

various interfaces such as a gearbox or bearing casings In one experiment consisting of

a pinion gear and an associated bearing a 44dB attenuation was seen between an AE

sensor placed directly on the pinion to one placed on the bearing casing and in some

cases intermediate loss of the signal was observed [20]

10

22 MEMS VIBRATION SENSORS

This thesis focuses on the development of a MEMS vibration sensor for fault

detection as the vibration signature of a mechanical system has the potential to give an

overview of the entire system Many types of MEMS vibration sensors exist with

different operating principals based on resonance the piezoelectric effect and

displacement variation

221 RESONANT SENSORS

Resonant sensors are based on the resonant frequency shift of a fixed-fixed micro-

beam that is excited into resonance using electrostatic thermal or piezoelectric methods

Under an applied axial load there is a shift in the resonant frequency which is expressed

as [21]

2 2

2

1 12 1

nn n

EI Fl2

fl A Eα γ

π ρ= +

I

(4)

where l is the beam length E is Youngrsquos Modulus I is the moment of inertia ρ is the

density A is the cross section area F is the applied axial force αn and γn are mode

dependant coefficients and are 4730 and 0295 for mode 1 in the case discussed The

axial force F is generated by a proof mass attached to one beam end which moves in

response to an external acceleration The performance of resonant sensors are dependant

on the quality factor Q which is the ratio of total energy stored in the system to the

energy lost per cycle and determines the sharpness and amplitude of the resonance peak

A high Q resonator develops a peak that is easily distinguishable in a phase locked loop

control circuit and allows for improved performance and resolution [3] A resonator Q

value is lowered by energy losses to the surrounding fluid attached supports and

11

material Losses to the fluid surrounding the MEMS structure is the dominant loss

mechanism for MEMS sensors operated in air and can be greatly reduced by vacuum

packaging which can be accomplished using glass-silicon anodic bonding [22] In

addition to increasing fabrication cost the long term stability of the vacuum degrades

with ultimate device failure stemming from leakage through micro cracks and defects

and out-gassing [22] Support losses arise from the restoring forces they generate and

can be reduce by balanced designs such as a double ended tuning fork (DETF) or

operation in higher modes [3] In addition the sensitivity of a resonant sensor can be

increased by the addition of micro-levers to increase the axial force In one such design a

two stage micro-lever increased the sensitivity by an order of magnitude over a more

conventional single lever arm [2324]

222 PIEZOELECTRIC SENSORS

Piezoelectric materials produce a charge in response to an applied force This is an

intrinsic effect in materials such as Zinc Oxide (ZnO) whereas materials such as Barium

Titanate or Lead Zirconate Titanate (PZT) need to be poled by placing the material in a

strong electric field at an elevated temperature [3] The piezoelectric effect is expressed

mathematical by [25]

i ij j ik kD d T Eε= + (5)

where D represents the electrical displacement (Cm2) d is the charge coefficient (CN)

T is the applied stress matrix (Nm2) ε is the permittivity matrix (Fm) and E is the

electric field matrix (Vm) The subscripts i=123 j=123456 and k=123 represent

the direction to which the physical property is related For micro-sensors poled in the out

of plane direction only the D3 term is relevant and if fabricated using ZnO or PZT only

12

the charge coefficients d31 d32 and d33 are relevant [25] The piezoelectric material is

commonly deposited as a thin film on the surface of compliant structures in bulk micro-

machined sensors or incorporated as a layer in surface micro-machining on a similar

structure [26] However the inclusion of multiple materials introduces large temperature

variations causing out of plane bending especially in surface micro-machined devices

[2627] Poled piezoelectric materials could be depolarized which serves to reduce the

piezoelectric affect in the material This can occur if the sensor is exposed to a strong

electric field of the opposite polarity high temperatures in excess of the Curie point or

high mechanical stress Also the piezoelectric effect reduces over time an effect that can

be reduced with the addition of composite elements [28] Foundry processes such as

PolyMUMPS and SoiMUMPS do not incorporate piezoelectric layers thus sensors using

this method of transduction must use a customized process The piezoelectric sensor

outlined in [26] uses polysilicon surface micro-machining similar to PolyMUMPS

however it incorporates a ZnO layer that is not available in the foundry process

223 DISPLACEMENT VARIATION SENSORS

Displacement variation sensors consist of a proof mass connected to a compliant

micro-flexure The movement of the proof mass is sensed using optical electron

tunneling or more commonly capacitive methods A recent optical accelerometer

utilized a novel multilayer nano-grating for in-plane displacement sensing of the proof

mass [29] For small space variations between the nano-gratings a large change in the

optical reflection amplitude of the grating was observed [30] Even though optical

sensors offer resolutions approaching the Brownian noise limit they are not available in

small packages that are easily placed in space constricted areas for a low cost [29]

13

The electron tunneling effect is observed when the proximity between two electrodes

one flexible and the other fixed is on the order of 10Ǻ [31] If the two electrodes are

biased with voltage VB a tunneling current It flows between the gap [31]

exp( )t B I tgI V xαinfin sdot minus Φ (6)

where α1 is a constant and equal to 1025 Ǻ-1 eV-frac12 Ф is the effective height of the

tunneling barrier and xtg is the tunneling gap When the flexible electrode moves in

response to acceleration a feedback controller maintains the original distance while

determining the applied acceleration Electron tunneling sensors are able to sense

accelerations in the microg-ng range due to the exponential relation between tunneling current

and gap However these highly sensitive devices are susceptible to thermo-mechanical

noise which could be reduced by operation at low pressure [32] It is also important for

the two electrodes to be coated with metal which is difficult to achieve with foundry

fabrication

Capacitive sensors are based on the parallel plate capacitor and are implemented

using inter-digitated comb fingers in an in-plane lateral or transverse drive or in an out of

plane parallel plate drive [33] These sensors have the advantage of high sensitivity low

power consumption small package low temperature dependence and are easily

integrated with existing foundry fabrication processes [34] The lateral drive offers a fully

linear response but low sensitivity whereas the transverse and parallel plate drives

achieve high sensitivity if gap distances are of a few microns and large capacitance areas

are utilized However both drives have a non-linear response and are prone to pull-in

Also if these sensors are fabricated using a surface micromachining process such as

PolyMUMPS where the ground plane is a few microns above the structural layer they

14

are prone to electrical failure due to shorting This was observed in an experiment where

similar sensors were placed in a vibrating environment with a peak of 120g and shorting

between the structural layer and ground plane was found to be the predominate mode of

failure [35] However with certain fabrication processes such as SoiMUMPS this issue

can be neglected as the substrate material below any moving component is removed

Most capacitance sensors are fabricated using an integrated CMOS (Complementary

Metal Oxide Semiconductor)-MEMS technique where the MEMS sensor is fabricated

before during or after the CMOS circuit fabrication [36] Although these sensors offer

high-sensitivity and small parasitic capacitance they suffer from in-plane and out of

plane curling of the beam which reduces the capacitance between adjacent comb fingers

[37] In addition multiple processing steps are required for device fabrication

More than the other transduction methods MEMS capacitive displacement sensors are

attractive for practical implementation in an automobile and therefore is the selected

method for the vibration sensor presented in this work Their low temperature dependence

is an ideal characteristic for an automobile environment that faces varying temperature If

implemented with off-chip readout circuitry they require a standard foundry fabrication

process which provides a strong platform for mass production and low unit cost

3 SENSOR STRUCTURE

This chapter presents the physical structure of the three MEMS vibration sensors

which includes the displacement sensing capacitive comb drive and the micro-

flexureproof mass system First the capacitive comb drive and the three variations

implemented the transverse drive lateral drive and differential tri-plate drive are

discussed Next the mechanical structure is presented and includes the micro-flexure

selection quality factor and equations of motion Finally instability that places a limit

on the operating range of the sensor as a result of a gap closing capacitive comb drive is

discussed

31 SENSING

Capacitive sensors are based on the parallel plate capacitor and are commonly

implemented using inter-digitated comb fingers A capacitance can be realized if a voltage

is placed between two closely spaced plates If the fringing electric field is neglected the

capacitance of the parallel plate capacitor Cpp (F) is expressed as [28]

air pppp

k AC

dε

= (7)

where kair=1 is the dielectric constant of air at atmospheric pressure ε=88542times10-12 F m-1

is the permittivity of free space App is the overlapping area (m2) and d is the distance

between the two plates (m)

15

16

Figure 3-1 An inter-digitated comb drive where changes in capacitance are generated by either changes in gap distance x0 or in the overlap area y0timest

Figure 3-1

For a set of inter-digitated comb fingers or comb drive shown in the total

capacitance Ctotal is expressed as [28]

0 0

0 0

( ) ( )( ) (Total cy y y yC N t

)x x x xε

+ += + + minus

(8)

where N is the number of comb fingers on the movable comb y0 is the initial comb finger

overlap length (m) tc is the out of plane comb finger thickness (m) and x0 is the inital

finger gap (m) In this work the comb drive is implemented in three arrangements

transverse lateral and differential tri-plate In a transverse comb drive changes in

capacitance are generated by transverse movements along the x-axis resulting in changes

to the gap x0 between the movable and fixed comb fingers The nonlinear relation

between total capacitance CTotalT and displacement is expressed as

17

00 0

1 1( ) (Total TC N ty

)x x x xε

= + minus +

(9)

Although the drive sensitivity part is not constant an average sensitivity can be

assumed for small displacements [38]

Total TC xpart

In a lateral comb drive changes in capacitance are generated by lateral movements

along the y-axis resulting in changes to the overlapping area y0timestc The total capacitance

CTotalL between all comb fingers is expressed as

00

2 ( )Total LN tC yx

yε= +

(10)

The sensitivity of the lateral comb drive SL is expressed as

0

2Total LL

C N tSy x

εpart= =

part

(11)

The lateral comb drive was first introduced by W C Tang T H Nguyen and R T

Howe [33] in an electrostatic resonator and has since been used numerous times in both

drive and sense configurations It has the advantage of a fully linear response and

negligible instability but offers low sensitivity The sensitivity for the lateral drive L1

implemented in this work is 0038 pF microm-1

A differential tri-plate drive shown in Figure 3-2 which is a variation of the

transverse drive is implemented to improve sensitivity and reduce noise [39]

18

Figure 3-2 An inter-digitated comb drive in a differential tri-plate arrangement where x01 x02 and x03 are the gap distances and x01= x02ltlt x03 to improve sensitivity

In the tri-plate drive the at rest gap distances between the two capacitors formed by any

movable finger is not equivalent For a differential arrangement insuring that x01=

x02ltltx03 the linearity for small displacements (less than 1 microm) is enhanced over the case

of a single ended arrangement where x01=x02=x03=x0 which is equivalent to the

transverse drive This is shown graphically in Figure 3-3 where x01= x02=x03=2 microm in

the lower curve and x01= x02= 2 microm and x03= 15 microm in the upper curve A substantial

improvement in linearity is attained with the tri-plate drive

19

0

005

01

015

02

025

03

035

0 02 04 06 08 1

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-3 Comparison of the effect of gap distances on linearity for two case lsquorsquo represents case 1 where the fingers in are implemented in a differential tri-plate arrangement and x01= x02= 2 microm and x03= 15 microm lsquorsquo represents the case 2 where the fingers are implemented in a single ended arrangement and x01= x02=x03=2 microm

The tri-plate drive is also implemented in a differential arrangement where the output

is the change in capacitance rather than the total capacitance resulting from an external

excitation This reduces errors cause by off-axis excitation and common mode noise The

output of the differential tri-plate drive is expressed as [39]

Total Top Total BottomC C C∆ = minus where

003 01

1 1( ) (Total TopC N ty

)x x x xε

= +

minus +

(12)

20

002 03

1 1( ) (Total BottomC N ty

)x x x xε

= +

minus +

A graphic comparison of the capacitancedisplacement relation for sensors T1 and T2 is

shown in Figure 3-4 For these gap closing drives it is assumed that for displacements

less that 1 microm the relation is linear indicated by the dashed line above each curve The

sensitivity within this region is approximately 0046 pF microm-1 and 0253 pF microm-1 for T1

and T2 respectively The order of magnitude improvement is attributed to the differential

tri-plate arrangement and thicker comb finger

0

02

04

06

08

1

12

14

16

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-4 lsquorsquo represents the output of the differential tri-plate drive while lsquorsquo represents the output of the transverse drive The dashed line represents the linear approximation for displacements less that 1 microm For the purpose of comparison the output of the transverse drive was equalized to zero at zero displacement

21

In this work the transverse drive is first implemented in sensor T1 using the

PolyMUMPS fabrication process Then in an effort to linearize the sensor response the

lateral drive is used in L1 using SoiMUMPS The reduction in drive sensitivity due to the

lateral drive is further reduced by increasing the gap distance to 3 microm which is

recommended by the foundry to ensure a successful device These reductions are

balanced by increasing the number of comb fingers N and selecting a substrate thickness

that is 125 times large than T1 The third device T2 is fabricated using SoiMUMPS and

uses the tri-plate differential drive This design capitalizes on the increase in sensitivity of

a gap closing drive and larger substrate thickness attaining an order of magnitude

increase in sensitivity over T1 and L1 The specifications for each drive along with the

successive improvement in drive sensitivity are outlined in Table 3-1

Table 3-1 Specifications for each capacitive comb drive implemented in this work T1 L1 and T2

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Comb Finger Thickness tc (microm) 2 25 25 Gap Distance x0 (microm) 2 3 2 Number of Fingers N 200 256 30 Overlap length y0 (microm) 45 10 70 Drive Sensitivity (pF microm-1) 0046 0038 0273

32 MECHANICAL ANALYSIS

321 MICRO-FLEXURE SELECTION

The sensor structure consists of a proof mass connected to a compliant micro-flexure

that itself is anchored to the sensor substrate The proof mass displaces in response to an

22

applied acceleration and this displacement is sensed using the capacitive comb drive

attached to the proof mass

Proof mass

Folded beam micro-flexure

Anchor

Figure 3-5 Sensor Structure

This section focuses on micro-flexure selection as it should exhibit a large linear

displacement range and have a high lateral stiffness ratio to reduce cross-axis sensitivity

A number of micro-flexures have been used in MEMS research these include the

clamped-clamped crab-leg and folded beam shown in Figure 3-6-through 3-7

The clamped-clamped flexure exhibits a high stiffness ratio however for large lateral

displacements extensional axial forces buildup within the beams resulting in a non-linear

force-displacement relation [40] Within the linear regime the stiffness of the clamped-

clamped flexure is expressed as [40]

3

3b

yEt wk

L=

(13)

23

where E is the Youngrsquos Modulus of the beam material (N m-2) tb is the beam out of

plane thickness (m) w is the beam width (m) and L is the beam length (m) The stiffness

ratio is expressed as [40]

2x

y

k Lk w

=

(14)

If L is two orders of magnitude greater than w this results in a ratio of approximately

10000 The linear displacement range is determined using Finite Element Analysis (FEA)

of a clamped-clamped flexure with L=387 microm w=4 microm and tb=2 microm The result of this

simulation is plotted in and compared to a structure with constant stiffness ky

= 035 N m-1 predicted by Eq (13) The linear relation is valid for small deflections that

are less than 2 microm In addition this micro-flexure is susceptible to buckling as a result of

a compressive residual stress that may arise due to the fabrication process [41]

Figure 3-9

In the crab leg design the thigh section is meant to relieve the extensional axial

stress postponing the occurrence of non-linearity until greater lateral displacement [42]

The stiffness of the crab leg flexure is expressed as [43]

3

44

b s s ty

s s t

Et w L LkL L L

κκ

+= +

(15)

where κ=IsIt=(wswt)3 is a ratio of the second moment of inertia of the shin (s) and thigh

(t) ws and wt represent the width and Ls and Lt represent the length of the shin and thigh

regions respectively The stiffness ratio is expressed as [43]

34

4t s s tx

y s t s

w L L Lkk w L L L

αα t

+= +

(16)

The linear displacement range of a crab leg flexure with Ls=285 microm Lt=30 microm ws=wt=3

microm and tb=2 microm is determined using FEA The results are plotted in Figure 3-9 and

compared to a structure of constant stiffness ky=035 N m-1 predicted by Eq (15) The

24

non-linearity is greatly reduced when compared to the clamp-clamped flexure with Eq

accurately predicting the deflection until 4 microm However the flexure is more

susceptible to off-axis deflection as there is an order of magnitude reduction in the

stiffness ratio as compared to the clamped-clamped flexure

(15)

The folded beam flexure stands out as it is designed to exhibit a large linear

displacement range and to relieve residual stress due to the manufacturing process by

allowing the beams to expand and contract along the axial direction [40] It was first used

in an electrostatic resonator and has subsequently been used for electrostatic actuation in

a magnetic hard drives and micro-positioning [3344] If the truss section is rigid

ItrussgtgtIbeam then the stiffness of the flexure is expressed as [4043]

3

32b

yEt wk

L=

(17)

The stiffness ratio is expressed as [40]

2x

y

k Lk w

=

(18)

Figure 3-6 Clamped-clamped flexure

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

AVERAGE EXPERIMENTAL SENSITIVITY WAS FOUND TO BE 0024 PF ΜM-1 FOR DISPLACEMENTS OF LESS THAN 1 ΜM 50

FIGURE 7-5 THE SOLID LINE REPRESENTS THE THEORETICAL CAPACITANCE TO VOLTAGE RELATION EXPRESSED BY EQ (37) WITH A SENSITIVITY OF 141 V PF-1 THE SOLID POINTS REPRESENT THE EXPERIMENTAL DATA POINTS WHEN DISCRETE CAPACITORS 1-4 PF ARE USED IN THE READOUT CIRCUIT THE HORIZONTAL DASHED LINE REPRESENTS THE OUTPUT VOLTAGE OF THE MEMS SENSOR AT REST AND CORRESPONDS TO A CAPACITANCE OF 35 PF 52

FIGURE 7-6 OSCILLOSCOPE TRACE OF THE INPUT EXCITATION VOLTAGE VE asymp 75 VPK AND OUTPUT OF THE VOLTAGE AMPLIFIER VVA asymp 103 VPK READINGS WERE TAKEN WITH THE MEMS SENSOR AT REST 53

FIGURE 7-7 OSCILLOSCOPE TRACE OF THE DEMODULATOR OUTPUT VDM asymp 76 VPK-PK AND OUTPUT OF THE LOW PASS FILTER VLP asymp 36 V READINGS WERE TAKEN WITH THE MEMS SENSOR AT REST 53

FIGURE 7-8 lsquorsquo REPRESENTS THE EXPERIMENTAL DATA POINTS WITH THE DASHED LINE REPRESENTING THE FITTED CURVE SOLID LINE REPRESENTS THE THEORETICAL DATA POINTS ACCORDING TO EQ (10) 54

FIGURE 7-9 EXPERIMENT SETUP WHICH INCLUDES THE MEMS SENOR AND READOUT BOARD MOUNTED ON THE CLAMPED-FREE BEAM AND TEST EQUIPMENT 55

FIGURE 7-10 CLOSE-UP OF THE MEMS SENSOR AND READOUT BOARD MOUNTED ON THE TEST BEAM ALONG WITH THE COMMERCIAL PIEZOELECTRIC ACCELEROMETER ONLY THE CHARGE AMPLIFIER (CA) AND THE VOLTAGE AMPLIFIER (VA) ARE MOUNTED ON THE SAME PROTOTYPE BOARD AS THE MEMS SENSOR DUE TO SPACE CONSTRAINTS 55

FIGURE 7-11 FIGURE COMPARING THE RESPONSE OF A COMMERCIAL PIEZOELECTRIC ACCELEROMETER AND MEMS SENSOR (L1) TO A BEAM VIBRATION A) DYNAMIC RESPONSE OF THE PIEZOELECTRIC ACCELEROMETER REPRESENTING THE TRUE BEAM VIBRATION WITH A PEAK VOLTAGE 114 V CORRESPONDING TO AN APPLIED ACCELERATION OF 114 G B) DYNAMIC RESPONSE OF THE MEMS SENSOR WITH A PEAK VOLTAGE OF 006 V 56

FIGURE 7-12 SENSOR DYNAMIC RESPONSE TO THE VIBRATION OF THE CLAMPED-FREE BEAM AND REPRESENTS A DIRECT EXPORT FROM THE DYNAMIC SIGNAL ANALYZER (AGILENT 35670) TOP TRACE IS THE TIME DOMAIN SIGNAL OF THE LOW PASS FILTER OUTPUT WHERE VLPMAXasymp73MV WHICH REPRESENTS A MAXIMUM ACCELERATION OF APPROXIMATELY 15G THE BOTTOM TRACE REPRESENTS THE FREQUENCY SPECTRUM WITH A PEAK AT 23HZ WHICH CORRESPONDS TO THE BEAM NATURAL FREQUENCY 57

FIGURE 7-13 EXPERIMENTAL RESULTS SHOWING THE RESPONSE OF THE SENSOR TO VARIOUS ACCELERATIONS IN THE RANGE OF 0-20 G WHERE lsquorsquo REPRESENTS THE EXPERIMENTAL DATA POINTS AND THE SOLID LINE REPRESENTS THE FITTED CURVE THE EXPERIMENTAL SENSITIVITY WAS FOUND TO BE 4258 MV G-1 58

vii

NOMENCLATURE a applied acceleration A applied acceleration amplitude App parallel plate overlapping area Asldie-film effective area for slide-film damping c coefficient of damping Cf feedback capacitance Cpp capacitance of the parallel plate capacitor cslide-film coefficient of damping as a result of slide-film damping

csqueeze-film coefficient of damping as a result of squeeze-film damping CTotal total capacitance of a comb drive CTotalBottom Total capacitance of the bottom drive in a tri-plate comb drive CTotalL total capacitance of a lateral drive CTotalT total capacitance of a transverse drive CTotalTop Total capacitance of the top drive in a tri-plate comb drive d distance between the parallel plates E Youngrsquos Modulus of sensor material eno noise voltage spectral density Eno total output noise voltage EnoAMP total output amplifier noise voltage EnoR total output resistor noise voltage

f excitation frequency in Hz fn system natural frequency in Hz

Fact electrostatic actuation force

FN noise force Fx-Elect total electrostatic force for pull-in instability Itruss second moment of inertia of truss section on the folded beam flexure Ibeam second moment of inertia of beam section on the folded beam

flexure kair dielectric constant of air at atmospheric pressure kB Boltzmannrsquos constant kT total stiffness of parallel folded beam flexure kx stiffness of various flexures in x-direction ky stiffness of various flexures in y-direction L flexure beam length Lplate effective plate length

viii

Ls shin beam length Lt thigh beam length m effective proof mass N number of comb fingers on the movable comb

NB number of balls

Qtotal total quality factor

Qf energy lost to surrounding fluid

Qs energy lost through supports

Qm energy lost through sensor material S theoretical sensitivity of sensor SL sensitivity of a lateral comb drive SM theoretical mechanical sensitivity of the proof massflexure system SRO sensitivity of readout circuit ST sensitivity of a transverse comb drive STAVG average sensitivity of a transverse comb drive t time T absolute temperature tb flexure beam thickness tc out of plane comb finger thickness

Vact electrostatic actuation voltage VCA output amplitude of charge amplifier VDM output amplitide of demodulator VE input amplitude of sinusoidal excitation voltage VLP output of low pass filter VLPMAX maximum output voltage at low pass filter VVA output amplitude of voltage amplifier vy relative motion of proof mass with respect to base Vy relative motion amplitude w flexure beam width Wplate effective plate width ws shin beam width wt thigh beam width x0 initial comb finger gap

x01 small gap distance on top side of the tri-plate comb drive

x02 small gap distance on bottom side of the tri-plate comb drive

ix

x03 larger gap distance on the tri-plate comb drive

XN noise displacement y0 initial comb finger overlap length yB base displacement YB base displacement amplitude ym absolute proof mass displacement α parameter dependant on the WplateLplate ratio ∆C Change in capacitance from a tri-plate drive in a differential

arrangement ε permittivity of free space κ second moment of inertia ratio (shinthigh) ζ damping ratio microair absolute viscosity of air ω base excitation frequency in rads ωn sensor natural frequency in rads ωv electrostatic actuation voltage frequency in rads L1 Sensor fabricated using SoiMUMPS with a lateral comb drive T1 Sensor fabricated using PolyMUMPS with a transverse comb drive T2 Sensor fabricated using SoiMUMPS with a tri-plate comb drive

x

1 INTRODUCTION

11 RESEARCH PURPOSE

The modern automobile consists of many mechanical systems such as power seats

windshield wipers mirrors trunks and windows which are all susceptible to breakdown

Without any condition monitoring system the breakdown is usually catastrophic and

requires an expensive part replacement Real-time condition monitoring allows for early

detection of faults which could require a simple solution such as the application of a

lubricant to fix This prolongs the useful life of the component and prevents sudden and

unexpected failure Real-time condition monitoring can be accomplished by examining the

vibration signature of a mechanical system For example an automobile power window

consists of a DC motor and its associated bearings and couplings a gear reduction system

consisting of worm and spur gears and kinematic links Faults resulting in excessive

vibrations may be caused by coupling misalignment bearing failure or gear train failure

Coupling misalignments occur at the connection between the drive shaft and the driven

shaft and are typically due to imperfect manufacturing Bearing failure is caused by lack

of lubrication or moisture contamination causing rusting while gear train failure is caused

by misaligned broken cracked or chipped gear teeth [1] Each fault occurs at a

characteristic frequency and so the state of the mechanical system can be determined by

monitoring the amplitudes of the relevant frequencies Vibrations due to coupling

misalignments occur at harmonics of the shaft rotational speed Gear vibrations occur at

the gear turn speed or at sidebands of the gear mesh frequency [1] Ball bearing vibrations

1

2

may be caused by outer bearing race defects fOB inner bearing race defects fIB ball

defects fB and train defects fT which all occur at specific frequencies [12]

cos( )12

dBOB r

p

bNf fd

β = minus

cos( )12

dBIB r

p

bNf fd

β = +

2cos( )1

2p r d

Bd p

d f bfb d

β = minus

cos( )12

drT

p

bffd

β = minus

(1)

where NB is the number of balls fr is the rotation frequency bd is the ball diameter β is the

contact angle between the ball and races and dp is the ball pitch diameter By monitoring

the real time vibration signature of the mechanical system anomalies can be quickly

identified and fixed

12 MICROELECTROMECHANICAL SENSORS

Microelectromechanical systems (MEMS) inertial sensors provide a small footprint

with sensitivities that are either comparable or exceed any macro sensor along with the

capability of mass production and low unit cost These sensors utilize compliant micro-

flexures attached to a proof mass that displaces in response to an environmental

acceleration Many transduction mechanisms have been developed that convert the

displacement into a measurable electric signal and include thermal piezoresistive

piezoelectric optical and capacitive methods Most MEMS sensors are silicon based and

are fabricated using surface or bulk micromachining Surface micromachining creates free

standing movable structures on top of a substrate using a combination of sacrificial layers

3

and structural layers which are commonly polysilicon [3] In bulk micromachining the

mechanical structures are defined using a removal process where bulk material typically

silicon is etched away MEMS sensors have been used for vibration and shock monitoring

on industrial systems and robotics guidance and navigation in global positioning systems

(GPS) seismometry in earthquake prediction image stabilization in digital cameras and

automobile safety and stability [4]

13 AUTOMOTIVE SENSORS

Sensors cover every major aspect of a modern automobile power-train sensors

monitor fuel combustion and emissions chassis sensors monitor road traction and tire

condition and body sensors facilitate air-bag deployment and vehicle proximity for radar

guided cruise control [5] Pressure sensors which typically consist of a piezoresistive

strain sensing element attached to a silicon diaphragm that deflects when exposed to an

applied pressure are one of the first micro-machined sensors used in an automobile

Implemented as manifold absolute pressure (MAP) sensors they allow precise control of

the air fuel ratio which allows the catalytic converter to efficiently reduce tailpipe

emissions [6] Variable reluctance sensors based on electromagnetics are used for

automobile traction control and produce a voltage output that is dependant on the

magnetic flux variations between a rotating component and the sensor bias magnet [57]

MEMS based linear accelerometers are utilized for airbag deployment upon impact and

provide a lower cost and smaller package over the older ball-in-tube accelerometers [6]

4

14 OPERATING ENVIRONMENT

Through proper packaging MEMS sensors can withstand the harsh automotive

environment where they face cyclic temperatures (-40degC to 150degC) mechanical shock or

vibration (10g-500g) and exposure to various fluids (oil brake fluid or engine coolant)

and airborne particles [8] Temperature changes induce thermal stresses causing

undesirable bending of the silicon microstructures that changes their dynamic response

Mechanical shock which could occur during a high impact collision or vibration from

uneven road conditions cause the delicate microstructures to fracture at points of stress

concentration These effects could be reduced by efficient design Use of a folded beam

micro-flexure that allows expansion and contraction or use of fillets at corners which

reduce the chance of fracture Bonding pads wire bonding and onboard circuit

components can be protected from air borne particles which could cause shorting using

parylene or gels Alternatively the entire package can be sealed hermetically which also

protects the MEMS device [8] In some cases such as fluid pressure sensors intimate

contact between the MEMS device and potentially corrosive fluid is necessary A recent

approach used a nanometer thin coating of silicon carbide (SiC) over a foundry fabricated

MEMS sensor using low temperature low pressure chemical vapor deposition (LPCVD)

[9] The SiC known for its chemical inertness did not compromise the mechanical

characteristics of the original design Therefore MEMS provides a strong platform for an

automotive sensor

5

15 OBJECTIVES AND CONTRIBUTION

The objective of this work is the development of a low frequency MEMS vibration

sensor The sensor must be low cost and mass producible which can be achieved using

standard foundry fabrication processes such as the Polysilicon Multi-User MEMS

Processes (PolyMUMPS) or the Silicon-on-Insulator Multi-User MEMS Processes

(SoiMUMPS) In addition the number of additional processing steps which are required

if integrated CMOS circuitry is used or if vacuum packaging is required must be

minimized This can be achieved using an off-chip circuit made with widely available

discrete components and sensor operation in air

In addition this work presents a MEMS sensor that is fabricated using standard

foundry processes such as PolyMUMPS and SoiMUMPS compared to many MEMS

sensors in literature that are fabricated using customized processes

16 RESEARCH OUTLINE

This work demonstrates three iterations of a low frequency capacitive vibration

sensor fabricated using the standard foundry processes Each sensor consists of a

suspended proof mass that displaces in response to an external vibration The proof mass

is attached to a folded beam micro-flexure compliant in one direction and its lateral

movement is sensed using a capacitive comb drive The first sensor T1 is fabricated in

PolyMUMPS and uses a transverse capacitive comb drive The second L1 fabricated in

SoiMUMPS uses a fully linear lateral capacitive drive and achieves an increase in

sensitivity from T1 Finally the third sensor T2 is fabricated in SoiMUMPS and uses a

differential tri-plate capacitive drive This sensor takes advantage of the increase in

6

sensitivity gained from a transverse drive but rejects off-axis displacement errors and

common mode using a differential tri-plate arrangement In addition an off-chip

capacitance to voltage readout circuit is fabricated and tested on sensor L1

Chapter 2 begins with an overview on the various methods of fault detection

followed by a review of MEMS vibration sensors each with a different transduction

mechanism Chapter 3 outlines the sensor structure focusing on the mechanical design

and sensing principal Chapter 4 focuses on the readout circuitry layout and simulation

results Chapter 5 outlines the various source of noise including mechanical noise arising

from the sensor and electric noise from the readout circuit Chapter 6 overviews the

foundry fabrication processes used including PolyMUMPs and SoiMUMPS The

experimental results of two sensors T1 and L1 which were fabricated and tested within

the thesis time frame are presented in Chapter 7 Chapter 8 concludes the thesis with a

brief summary and outline for future work

2 LITERATURE REVIEW

21 EXISTING METHODS FOR FAULT DETECTION

As this thesis develops a sensor for fault detection using vibration monitoring this

section overviews three existing methods for fault detection in mechanical systems First

motor current monitoring is discussed as it has the ability to measure significant bearing

faults in motor systems Next temperature sensors specifically thin film thermocouples

(TFTC) and MEMS temperature sensors are discussed These sensors are mounted in the

chamber of a lubricating fluid or in close proximity to a moving component and monitor

temperature fluctuations caused by increased wear Finally acoustic emission (AE)

sensors are discussed which have the ability to detect ultrasonic emissions generated by

plastic deformation in the crystalline lattice of a mechanical component

211 MOTOR CURRENT MONITORING

Monitoring of stator current is used as a method of determining bearing related faults

that produce radial motion between rotor and stator This motion varies the air gap flux

density producing stator currents at predictable frequencies fbng that represent the

bearing faults [10]

bng e vf f m f= plusmn sdot (2)

where fe is the electrical supply frequency m=123hellip and fv is one of the characteristics

frequencies listed in the equations of (1)The bearing damage must be significant to

produce noticeable current fluctuations and it may be difficult to distinguish from current

fluctuations due to voltage imbalances machine saturation or statorarmature faults [2]

7

8

212 TEMPERATURE SENSORS

Temperature sensors are largely used in high temperature high load and high speed

environments typically turbine engines and industrial tools Thermocouple sensors

consist of two dissimilar metals joined together at two junctions and when exposed to a

temperature difference generate a thermoelectric voltage dependant only on the material

properties and junction temperature difference [11] In turbine engines these sensors can

be placed on bearing casings or in the lubricant to monitor temperature fluctuations due

to increased friction The thermocouples are fabricated using bulk materials or micro-

machined thin films (TFTC) with the latter providing a faster response small package

and substantially lower production costs [12] The TFTC can be embedded for in situ

monitoring of processes in hostile environments One study successfully embedded

TFTC into electroplated nickel and attained device characteristics on par with surface

mounted devices [13] More recently MEMS temperature sensors have been developed

that consist of micro-machined semiconductor material that undergoes structural

deformation with temperature changes A common implementation is a bent beam

structure whose temperature induced deflection δ is expressed as [14]

( ) 3

2 2

sin12 cos sinz

EA T LEI AEL

α βδ 2β β

∆=

+

(3)

Temperature sensors are well suited for localized fault monitoring and detection but can

only characterize a mechanical system consisting of multiple components through the use

of multiple sensors

9

213 ACOUSTIC EMISSION SENSORS

Acoustic emission (AE) sensors have been used to characterize wear in machine

tools and monitor bearing and gear problems in centrifugal pumps [1516] First

developed as a Non-Destructive Testing (NDT) technique to detect cracks in civil

structures these sensors detect acoustic emissions generated by the release of vibration

waves in a crystalline lattice due to plastic deformation or crack growth [16]

Measurements are made using piezoelectric transducers with high natural frequencies

100 kHz to 1 MHz to capture the ultrasonic AE emissions An AE sensor is useful as it

has the ability to detect subsurface cracks in gear teeth or bearings before appearing on

the surface causing further damage [17] More recently MEMS acoustic sensors have

been developed and one design includes multiple transducers on a single substrate

which each detect acoustic emission energy at different frequencies [18] This helps

distinguish spurious acoustic emissions arising from impact and friction from those

arising from plastic deformation When compared to typical piezoelectric sensors the

MEMS devices have lower sensitivities and fail to detect some acoustic emissions

[1819] In addition the acoustic emission signal suffers severe attenuation as is crosses

various interfaces such as a gearbox or bearing casings In one experiment consisting of

a pinion gear and an associated bearing a 44dB attenuation was seen between an AE

sensor placed directly on the pinion to one placed on the bearing casing and in some

cases intermediate loss of the signal was observed [20]

10

22 MEMS VIBRATION SENSORS

This thesis focuses on the development of a MEMS vibration sensor for fault

detection as the vibration signature of a mechanical system has the potential to give an

overview of the entire system Many types of MEMS vibration sensors exist with

different operating principals based on resonance the piezoelectric effect and

displacement variation

221 RESONANT SENSORS

Resonant sensors are based on the resonant frequency shift of a fixed-fixed micro-

beam that is excited into resonance using electrostatic thermal or piezoelectric methods

Under an applied axial load there is a shift in the resonant frequency which is expressed

as [21]

2 2

2

1 12 1

nn n

EI Fl2

fl A Eα γ

π ρ= +

I

(4)

where l is the beam length E is Youngrsquos Modulus I is the moment of inertia ρ is the

density A is the cross section area F is the applied axial force αn and γn are mode

dependant coefficients and are 4730 and 0295 for mode 1 in the case discussed The

axial force F is generated by a proof mass attached to one beam end which moves in

response to an external acceleration The performance of resonant sensors are dependant

on the quality factor Q which is the ratio of total energy stored in the system to the

energy lost per cycle and determines the sharpness and amplitude of the resonance peak

A high Q resonator develops a peak that is easily distinguishable in a phase locked loop

control circuit and allows for improved performance and resolution [3] A resonator Q

value is lowered by energy losses to the surrounding fluid attached supports and

11

material Losses to the fluid surrounding the MEMS structure is the dominant loss

mechanism for MEMS sensors operated in air and can be greatly reduced by vacuum

packaging which can be accomplished using glass-silicon anodic bonding [22] In

addition to increasing fabrication cost the long term stability of the vacuum degrades

with ultimate device failure stemming from leakage through micro cracks and defects

and out-gassing [22] Support losses arise from the restoring forces they generate and

can be reduce by balanced designs such as a double ended tuning fork (DETF) or

operation in higher modes [3] In addition the sensitivity of a resonant sensor can be

increased by the addition of micro-levers to increase the axial force In one such design a

two stage micro-lever increased the sensitivity by an order of magnitude over a more

conventional single lever arm [2324]

222 PIEZOELECTRIC SENSORS

Piezoelectric materials produce a charge in response to an applied force This is an

intrinsic effect in materials such as Zinc Oxide (ZnO) whereas materials such as Barium

Titanate or Lead Zirconate Titanate (PZT) need to be poled by placing the material in a

strong electric field at an elevated temperature [3] The piezoelectric effect is expressed

mathematical by [25]

i ij j ik kD d T Eε= + (5)

where D represents the electrical displacement (Cm2) d is the charge coefficient (CN)

T is the applied stress matrix (Nm2) ε is the permittivity matrix (Fm) and E is the

electric field matrix (Vm) The subscripts i=123 j=123456 and k=123 represent

the direction to which the physical property is related For micro-sensors poled in the out

of plane direction only the D3 term is relevant and if fabricated using ZnO or PZT only

12

the charge coefficients d31 d32 and d33 are relevant [25] The piezoelectric material is

commonly deposited as a thin film on the surface of compliant structures in bulk micro-

machined sensors or incorporated as a layer in surface micro-machining on a similar

structure [26] However the inclusion of multiple materials introduces large temperature

variations causing out of plane bending especially in surface micro-machined devices

[2627] Poled piezoelectric materials could be depolarized which serves to reduce the

piezoelectric affect in the material This can occur if the sensor is exposed to a strong

electric field of the opposite polarity high temperatures in excess of the Curie point or

high mechanical stress Also the piezoelectric effect reduces over time an effect that can

be reduced with the addition of composite elements [28] Foundry processes such as

PolyMUMPS and SoiMUMPS do not incorporate piezoelectric layers thus sensors using

this method of transduction must use a customized process The piezoelectric sensor

outlined in [26] uses polysilicon surface micro-machining similar to PolyMUMPS

however it incorporates a ZnO layer that is not available in the foundry process

223 DISPLACEMENT VARIATION SENSORS

Displacement variation sensors consist of a proof mass connected to a compliant

micro-flexure The movement of the proof mass is sensed using optical electron

tunneling or more commonly capacitive methods A recent optical accelerometer

utilized a novel multilayer nano-grating for in-plane displacement sensing of the proof

mass [29] For small space variations between the nano-gratings a large change in the

optical reflection amplitude of the grating was observed [30] Even though optical

sensors offer resolutions approaching the Brownian noise limit they are not available in

small packages that are easily placed in space constricted areas for a low cost [29]

13

The electron tunneling effect is observed when the proximity between two electrodes

one flexible and the other fixed is on the order of 10Ǻ [31] If the two electrodes are

biased with voltage VB a tunneling current It flows between the gap [31]

exp( )t B I tgI V xαinfin sdot minus Φ (6)

where α1 is a constant and equal to 1025 Ǻ-1 eV-frac12 Ф is the effective height of the

tunneling barrier and xtg is the tunneling gap When the flexible electrode moves in

response to acceleration a feedback controller maintains the original distance while

determining the applied acceleration Electron tunneling sensors are able to sense

accelerations in the microg-ng range due to the exponential relation between tunneling current

and gap However these highly sensitive devices are susceptible to thermo-mechanical

noise which could be reduced by operation at low pressure [32] It is also important for

the two electrodes to be coated with metal which is difficult to achieve with foundry

fabrication

Capacitive sensors are based on the parallel plate capacitor and are implemented

using inter-digitated comb fingers in an in-plane lateral or transverse drive or in an out of

plane parallel plate drive [33] These sensors have the advantage of high sensitivity low

power consumption small package low temperature dependence and are easily

integrated with existing foundry fabrication processes [34] The lateral drive offers a fully

linear response but low sensitivity whereas the transverse and parallel plate drives

achieve high sensitivity if gap distances are of a few microns and large capacitance areas

are utilized However both drives have a non-linear response and are prone to pull-in

Also if these sensors are fabricated using a surface micromachining process such as

PolyMUMPS where the ground plane is a few microns above the structural layer they

14

are prone to electrical failure due to shorting This was observed in an experiment where

similar sensors were placed in a vibrating environment with a peak of 120g and shorting

between the structural layer and ground plane was found to be the predominate mode of

failure [35] However with certain fabrication processes such as SoiMUMPS this issue

can be neglected as the substrate material below any moving component is removed

Most capacitance sensors are fabricated using an integrated CMOS (Complementary

Metal Oxide Semiconductor)-MEMS technique where the MEMS sensor is fabricated

before during or after the CMOS circuit fabrication [36] Although these sensors offer

high-sensitivity and small parasitic capacitance they suffer from in-plane and out of

plane curling of the beam which reduces the capacitance between adjacent comb fingers

[37] In addition multiple processing steps are required for device fabrication

More than the other transduction methods MEMS capacitive displacement sensors are

attractive for practical implementation in an automobile and therefore is the selected

method for the vibration sensor presented in this work Their low temperature dependence

is an ideal characteristic for an automobile environment that faces varying temperature If

implemented with off-chip readout circuitry they require a standard foundry fabrication

process which provides a strong platform for mass production and low unit cost

3 SENSOR STRUCTURE

This chapter presents the physical structure of the three MEMS vibration sensors

which includes the displacement sensing capacitive comb drive and the micro-

flexureproof mass system First the capacitive comb drive and the three variations

implemented the transverse drive lateral drive and differential tri-plate drive are

discussed Next the mechanical structure is presented and includes the micro-flexure

selection quality factor and equations of motion Finally instability that places a limit

on the operating range of the sensor as a result of a gap closing capacitive comb drive is

discussed

31 SENSING

Capacitive sensors are based on the parallel plate capacitor and are commonly

implemented using inter-digitated comb fingers A capacitance can be realized if a voltage

is placed between two closely spaced plates If the fringing electric field is neglected the

capacitance of the parallel plate capacitor Cpp (F) is expressed as [28]

air pppp

k AC

dε

= (7)

where kair=1 is the dielectric constant of air at atmospheric pressure ε=88542times10-12 F m-1

is the permittivity of free space App is the overlapping area (m2) and d is the distance

between the two plates (m)

15

16

Figure 3-1 An inter-digitated comb drive where changes in capacitance are generated by either changes in gap distance x0 or in the overlap area y0timest

Figure 3-1

For a set of inter-digitated comb fingers or comb drive shown in the total

capacitance Ctotal is expressed as [28]

0 0

0 0

( ) ( )( ) (Total cy y y yC N t

)x x x xε

+ += + + minus

(8)

where N is the number of comb fingers on the movable comb y0 is the initial comb finger

overlap length (m) tc is the out of plane comb finger thickness (m) and x0 is the inital

finger gap (m) In this work the comb drive is implemented in three arrangements

transverse lateral and differential tri-plate In a transverse comb drive changes in

capacitance are generated by transverse movements along the x-axis resulting in changes

to the gap x0 between the movable and fixed comb fingers The nonlinear relation

between total capacitance CTotalT and displacement is expressed as

17

00 0

1 1( ) (Total TC N ty

)x x x xε

= + minus +

(9)

Although the drive sensitivity part is not constant an average sensitivity can be

assumed for small displacements [38]

Total TC xpart

In a lateral comb drive changes in capacitance are generated by lateral movements

along the y-axis resulting in changes to the overlapping area y0timestc The total capacitance

CTotalL between all comb fingers is expressed as

00

2 ( )Total LN tC yx

yε= +

(10)

The sensitivity of the lateral comb drive SL is expressed as

0

2Total LL

C N tSy x

εpart= =

part

(11)

The lateral comb drive was first introduced by W C Tang T H Nguyen and R T

Howe [33] in an electrostatic resonator and has since been used numerous times in both

drive and sense configurations It has the advantage of a fully linear response and

negligible instability but offers low sensitivity The sensitivity for the lateral drive L1

implemented in this work is 0038 pF microm-1

A differential tri-plate drive shown in Figure 3-2 which is a variation of the

transverse drive is implemented to improve sensitivity and reduce noise [39]

18

Figure 3-2 An inter-digitated comb drive in a differential tri-plate arrangement where x01 x02 and x03 are the gap distances and x01= x02ltlt x03 to improve sensitivity

In the tri-plate drive the at rest gap distances between the two capacitors formed by any

movable finger is not equivalent For a differential arrangement insuring that x01=

x02ltltx03 the linearity for small displacements (less than 1 microm) is enhanced over the case

of a single ended arrangement where x01=x02=x03=x0 which is equivalent to the

transverse drive This is shown graphically in Figure 3-3 where x01= x02=x03=2 microm in

the lower curve and x01= x02= 2 microm and x03= 15 microm in the upper curve A substantial

improvement in linearity is attained with the tri-plate drive

19

0

005

01

015

02

025

03

035

0 02 04 06 08 1

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-3 Comparison of the effect of gap distances on linearity for two case lsquorsquo represents case 1 where the fingers in are implemented in a differential tri-plate arrangement and x01= x02= 2 microm and x03= 15 microm lsquorsquo represents the case 2 where the fingers are implemented in a single ended arrangement and x01= x02=x03=2 microm

The tri-plate drive is also implemented in a differential arrangement where the output

is the change in capacitance rather than the total capacitance resulting from an external

excitation This reduces errors cause by off-axis excitation and common mode noise The

output of the differential tri-plate drive is expressed as [39]

Total Top Total BottomC C C∆ = minus where

003 01

1 1( ) (Total TopC N ty

)x x x xε

= +

minus +

(12)

20

002 03

1 1( ) (Total BottomC N ty

)x x x xε

= +

minus +

A graphic comparison of the capacitancedisplacement relation for sensors T1 and T2 is

shown in Figure 3-4 For these gap closing drives it is assumed that for displacements

less that 1 microm the relation is linear indicated by the dashed line above each curve The

sensitivity within this region is approximately 0046 pF microm-1 and 0253 pF microm-1 for T1

and T2 respectively The order of magnitude improvement is attributed to the differential

tri-plate arrangement and thicker comb finger

0

02

04

06

08

1

12

14

16

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-4 lsquorsquo represents the output of the differential tri-plate drive while lsquorsquo represents the output of the transverse drive The dashed line represents the linear approximation for displacements less that 1 microm For the purpose of comparison the output of the transverse drive was equalized to zero at zero displacement

21

In this work the transverse drive is first implemented in sensor T1 using the

PolyMUMPS fabrication process Then in an effort to linearize the sensor response the

lateral drive is used in L1 using SoiMUMPS The reduction in drive sensitivity due to the

lateral drive is further reduced by increasing the gap distance to 3 microm which is

recommended by the foundry to ensure a successful device These reductions are

balanced by increasing the number of comb fingers N and selecting a substrate thickness

that is 125 times large than T1 The third device T2 is fabricated using SoiMUMPS and

uses the tri-plate differential drive This design capitalizes on the increase in sensitivity of

a gap closing drive and larger substrate thickness attaining an order of magnitude

increase in sensitivity over T1 and L1 The specifications for each drive along with the

successive improvement in drive sensitivity are outlined in Table 3-1

Table 3-1 Specifications for each capacitive comb drive implemented in this work T1 L1 and T2

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Comb Finger Thickness tc (microm) 2 25 25 Gap Distance x0 (microm) 2 3 2 Number of Fingers N 200 256 30 Overlap length y0 (microm) 45 10 70 Drive Sensitivity (pF microm-1) 0046 0038 0273

32 MECHANICAL ANALYSIS

321 MICRO-FLEXURE SELECTION

The sensor structure consists of a proof mass connected to a compliant micro-flexure

that itself is anchored to the sensor substrate The proof mass displaces in response to an

22

applied acceleration and this displacement is sensed using the capacitive comb drive

attached to the proof mass

Proof mass

Folded beam micro-flexure

Anchor

Figure 3-5 Sensor Structure

This section focuses on micro-flexure selection as it should exhibit a large linear

displacement range and have a high lateral stiffness ratio to reduce cross-axis sensitivity

A number of micro-flexures have been used in MEMS research these include the

clamped-clamped crab-leg and folded beam shown in Figure 3-6-through 3-7

The clamped-clamped flexure exhibits a high stiffness ratio however for large lateral

displacements extensional axial forces buildup within the beams resulting in a non-linear

force-displacement relation [40] Within the linear regime the stiffness of the clamped-

clamped flexure is expressed as [40]

3

3b

yEt wk

L=

(13)

23

where E is the Youngrsquos Modulus of the beam material (N m-2) tb is the beam out of

plane thickness (m) w is the beam width (m) and L is the beam length (m) The stiffness

ratio is expressed as [40]

2x

y

k Lk w

=

(14)

If L is two orders of magnitude greater than w this results in a ratio of approximately

10000 The linear displacement range is determined using Finite Element Analysis (FEA)

of a clamped-clamped flexure with L=387 microm w=4 microm and tb=2 microm The result of this

simulation is plotted in and compared to a structure with constant stiffness ky

= 035 N m-1 predicted by Eq (13) The linear relation is valid for small deflections that

are less than 2 microm In addition this micro-flexure is susceptible to buckling as a result of

a compressive residual stress that may arise due to the fabrication process [41]

Figure 3-9

In the crab leg design the thigh section is meant to relieve the extensional axial

stress postponing the occurrence of non-linearity until greater lateral displacement [42]

The stiffness of the crab leg flexure is expressed as [43]

3

44

b s s ty

s s t

Et w L LkL L L

κκ

+= +

(15)

where κ=IsIt=(wswt)3 is a ratio of the second moment of inertia of the shin (s) and thigh

(t) ws and wt represent the width and Ls and Lt represent the length of the shin and thigh

regions respectively The stiffness ratio is expressed as [43]

34

4t s s tx

y s t s

w L L Lkk w L L L

αα t

+= +

(16)

The linear displacement range of a crab leg flexure with Ls=285 microm Lt=30 microm ws=wt=3

microm and tb=2 microm is determined using FEA The results are plotted in Figure 3-9 and

compared to a structure of constant stiffness ky=035 N m-1 predicted by Eq (15) The

24

non-linearity is greatly reduced when compared to the clamp-clamped flexure with Eq

accurately predicting the deflection until 4 microm However the flexure is more

susceptible to off-axis deflection as there is an order of magnitude reduction in the

stiffness ratio as compared to the clamped-clamped flexure

(15)

The folded beam flexure stands out as it is designed to exhibit a large linear

displacement range and to relieve residual stress due to the manufacturing process by

allowing the beams to expand and contract along the axial direction [40] It was first used

in an electrostatic resonator and has subsequently been used for electrostatic actuation in

a magnetic hard drives and micro-positioning [3344] If the truss section is rigid

ItrussgtgtIbeam then the stiffness of the flexure is expressed as [4043]

3

32b

yEt wk

L=

(17)

The stiffness ratio is expressed as [40]

2x

y

k Lk w

=

(18)

Figure 3-6 Clamped-clamped flexure

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

NOMENCLATURE a applied acceleration A applied acceleration amplitude App parallel plate overlapping area Asldie-film effective area for slide-film damping c coefficient of damping Cf feedback capacitance Cpp capacitance of the parallel plate capacitor cslide-film coefficient of damping as a result of slide-film damping

csqueeze-film coefficient of damping as a result of squeeze-film damping CTotal total capacitance of a comb drive CTotalBottom Total capacitance of the bottom drive in a tri-plate comb drive CTotalL total capacitance of a lateral drive CTotalT total capacitance of a transverse drive CTotalTop Total capacitance of the top drive in a tri-plate comb drive d distance between the parallel plates E Youngrsquos Modulus of sensor material eno noise voltage spectral density Eno total output noise voltage EnoAMP total output amplifier noise voltage EnoR total output resistor noise voltage

f excitation frequency in Hz fn system natural frequency in Hz

Fact electrostatic actuation force

FN noise force Fx-Elect total electrostatic force for pull-in instability Itruss second moment of inertia of truss section on the folded beam flexure Ibeam second moment of inertia of beam section on the folded beam

flexure kair dielectric constant of air at atmospheric pressure kB Boltzmannrsquos constant kT total stiffness of parallel folded beam flexure kx stiffness of various flexures in x-direction ky stiffness of various flexures in y-direction L flexure beam length Lplate effective plate length

viii

Ls shin beam length Lt thigh beam length m effective proof mass N number of comb fingers on the movable comb

NB number of balls

Qtotal total quality factor

Qf energy lost to surrounding fluid

Qs energy lost through supports

Qm energy lost through sensor material S theoretical sensitivity of sensor SL sensitivity of a lateral comb drive SM theoretical mechanical sensitivity of the proof massflexure system SRO sensitivity of readout circuit ST sensitivity of a transverse comb drive STAVG average sensitivity of a transverse comb drive t time T absolute temperature tb flexure beam thickness tc out of plane comb finger thickness

Vact electrostatic actuation voltage VCA output amplitude of charge amplifier VDM output amplitide of demodulator VE input amplitude of sinusoidal excitation voltage VLP output of low pass filter VLPMAX maximum output voltage at low pass filter VVA output amplitude of voltage amplifier vy relative motion of proof mass with respect to base Vy relative motion amplitude w flexure beam width Wplate effective plate width ws shin beam width wt thigh beam width x0 initial comb finger gap

x01 small gap distance on top side of the tri-plate comb drive

x02 small gap distance on bottom side of the tri-plate comb drive

ix

x03 larger gap distance on the tri-plate comb drive

XN noise displacement y0 initial comb finger overlap length yB base displacement YB base displacement amplitude ym absolute proof mass displacement α parameter dependant on the WplateLplate ratio ∆C Change in capacitance from a tri-plate drive in a differential

arrangement ε permittivity of free space κ second moment of inertia ratio (shinthigh) ζ damping ratio microair absolute viscosity of air ω base excitation frequency in rads ωn sensor natural frequency in rads ωv electrostatic actuation voltage frequency in rads L1 Sensor fabricated using SoiMUMPS with a lateral comb drive T1 Sensor fabricated using PolyMUMPS with a transverse comb drive T2 Sensor fabricated using SoiMUMPS with a tri-plate comb drive

x

1 INTRODUCTION

11 RESEARCH PURPOSE

The modern automobile consists of many mechanical systems such as power seats

windshield wipers mirrors trunks and windows which are all susceptible to breakdown

Without any condition monitoring system the breakdown is usually catastrophic and

requires an expensive part replacement Real-time condition monitoring allows for early

detection of faults which could require a simple solution such as the application of a

lubricant to fix This prolongs the useful life of the component and prevents sudden and

unexpected failure Real-time condition monitoring can be accomplished by examining the

vibration signature of a mechanical system For example an automobile power window

consists of a DC motor and its associated bearings and couplings a gear reduction system

consisting of worm and spur gears and kinematic links Faults resulting in excessive

vibrations may be caused by coupling misalignment bearing failure or gear train failure

Coupling misalignments occur at the connection between the drive shaft and the driven

shaft and are typically due to imperfect manufacturing Bearing failure is caused by lack

of lubrication or moisture contamination causing rusting while gear train failure is caused

by misaligned broken cracked or chipped gear teeth [1] Each fault occurs at a

characteristic frequency and so the state of the mechanical system can be determined by

monitoring the amplitudes of the relevant frequencies Vibrations due to coupling

misalignments occur at harmonics of the shaft rotational speed Gear vibrations occur at

the gear turn speed or at sidebands of the gear mesh frequency [1] Ball bearing vibrations

1

2

may be caused by outer bearing race defects fOB inner bearing race defects fIB ball

defects fB and train defects fT which all occur at specific frequencies [12]

cos( )12

dBOB r

p

bNf fd

β = minus

cos( )12

dBIB r

p

bNf fd

β = +

2cos( )1

2p r d

Bd p

d f bfb d

β = minus

cos( )12

drT

p

bffd

β = minus

(1)

where NB is the number of balls fr is the rotation frequency bd is the ball diameter β is the

contact angle between the ball and races and dp is the ball pitch diameter By monitoring

the real time vibration signature of the mechanical system anomalies can be quickly

identified and fixed

12 MICROELECTROMECHANICAL SENSORS

Microelectromechanical systems (MEMS) inertial sensors provide a small footprint

with sensitivities that are either comparable or exceed any macro sensor along with the

capability of mass production and low unit cost These sensors utilize compliant micro-

flexures attached to a proof mass that displaces in response to an environmental

acceleration Many transduction mechanisms have been developed that convert the

displacement into a measurable electric signal and include thermal piezoresistive

piezoelectric optical and capacitive methods Most MEMS sensors are silicon based and

are fabricated using surface or bulk micromachining Surface micromachining creates free

standing movable structures on top of a substrate using a combination of sacrificial layers

3

and structural layers which are commonly polysilicon [3] In bulk micromachining the

mechanical structures are defined using a removal process where bulk material typically

silicon is etched away MEMS sensors have been used for vibration and shock monitoring

on industrial systems and robotics guidance and navigation in global positioning systems

(GPS) seismometry in earthquake prediction image stabilization in digital cameras and

automobile safety and stability [4]

13 AUTOMOTIVE SENSORS

Sensors cover every major aspect of a modern automobile power-train sensors

monitor fuel combustion and emissions chassis sensors monitor road traction and tire

condition and body sensors facilitate air-bag deployment and vehicle proximity for radar

guided cruise control [5] Pressure sensors which typically consist of a piezoresistive

strain sensing element attached to a silicon diaphragm that deflects when exposed to an

applied pressure are one of the first micro-machined sensors used in an automobile

Implemented as manifold absolute pressure (MAP) sensors they allow precise control of

the air fuel ratio which allows the catalytic converter to efficiently reduce tailpipe

emissions [6] Variable reluctance sensors based on electromagnetics are used for

automobile traction control and produce a voltage output that is dependant on the

magnetic flux variations between a rotating component and the sensor bias magnet [57]

MEMS based linear accelerometers are utilized for airbag deployment upon impact and

provide a lower cost and smaller package over the older ball-in-tube accelerometers [6]

4

14 OPERATING ENVIRONMENT

Through proper packaging MEMS sensors can withstand the harsh automotive

environment where they face cyclic temperatures (-40degC to 150degC) mechanical shock or

vibration (10g-500g) and exposure to various fluids (oil brake fluid or engine coolant)

and airborne particles [8] Temperature changes induce thermal stresses causing

undesirable bending of the silicon microstructures that changes their dynamic response

Mechanical shock which could occur during a high impact collision or vibration from

uneven road conditions cause the delicate microstructures to fracture at points of stress

concentration These effects could be reduced by efficient design Use of a folded beam

micro-flexure that allows expansion and contraction or use of fillets at corners which

reduce the chance of fracture Bonding pads wire bonding and onboard circuit

components can be protected from air borne particles which could cause shorting using

parylene or gels Alternatively the entire package can be sealed hermetically which also

protects the MEMS device [8] In some cases such as fluid pressure sensors intimate

contact between the MEMS device and potentially corrosive fluid is necessary A recent

approach used a nanometer thin coating of silicon carbide (SiC) over a foundry fabricated

MEMS sensor using low temperature low pressure chemical vapor deposition (LPCVD)

[9] The SiC known for its chemical inertness did not compromise the mechanical

characteristics of the original design Therefore MEMS provides a strong platform for an

automotive sensor

5

15 OBJECTIVES AND CONTRIBUTION

The objective of this work is the development of a low frequency MEMS vibration

sensor The sensor must be low cost and mass producible which can be achieved using

standard foundry fabrication processes such as the Polysilicon Multi-User MEMS

Processes (PolyMUMPS) or the Silicon-on-Insulator Multi-User MEMS Processes

(SoiMUMPS) In addition the number of additional processing steps which are required

if integrated CMOS circuitry is used or if vacuum packaging is required must be

minimized This can be achieved using an off-chip circuit made with widely available

discrete components and sensor operation in air

In addition this work presents a MEMS sensor that is fabricated using standard

foundry processes such as PolyMUMPS and SoiMUMPS compared to many MEMS

sensors in literature that are fabricated using customized processes

16 RESEARCH OUTLINE

This work demonstrates three iterations of a low frequency capacitive vibration

sensor fabricated using the standard foundry processes Each sensor consists of a

suspended proof mass that displaces in response to an external vibration The proof mass

is attached to a folded beam micro-flexure compliant in one direction and its lateral

movement is sensed using a capacitive comb drive The first sensor T1 is fabricated in

PolyMUMPS and uses a transverse capacitive comb drive The second L1 fabricated in

SoiMUMPS uses a fully linear lateral capacitive drive and achieves an increase in

sensitivity from T1 Finally the third sensor T2 is fabricated in SoiMUMPS and uses a

differential tri-plate capacitive drive This sensor takes advantage of the increase in

6

sensitivity gained from a transverse drive but rejects off-axis displacement errors and

common mode using a differential tri-plate arrangement In addition an off-chip

capacitance to voltage readout circuit is fabricated and tested on sensor L1

Chapter 2 begins with an overview on the various methods of fault detection

followed by a review of MEMS vibration sensors each with a different transduction

mechanism Chapter 3 outlines the sensor structure focusing on the mechanical design

and sensing principal Chapter 4 focuses on the readout circuitry layout and simulation

results Chapter 5 outlines the various source of noise including mechanical noise arising

from the sensor and electric noise from the readout circuit Chapter 6 overviews the

foundry fabrication processes used including PolyMUMPs and SoiMUMPS The

experimental results of two sensors T1 and L1 which were fabricated and tested within

the thesis time frame are presented in Chapter 7 Chapter 8 concludes the thesis with a

brief summary and outline for future work

2 LITERATURE REVIEW

21 EXISTING METHODS FOR FAULT DETECTION

As this thesis develops a sensor for fault detection using vibration monitoring this

section overviews three existing methods for fault detection in mechanical systems First

motor current monitoring is discussed as it has the ability to measure significant bearing

faults in motor systems Next temperature sensors specifically thin film thermocouples

(TFTC) and MEMS temperature sensors are discussed These sensors are mounted in the

chamber of a lubricating fluid or in close proximity to a moving component and monitor

temperature fluctuations caused by increased wear Finally acoustic emission (AE)

sensors are discussed which have the ability to detect ultrasonic emissions generated by

plastic deformation in the crystalline lattice of a mechanical component

211 MOTOR CURRENT MONITORING

Monitoring of stator current is used as a method of determining bearing related faults

that produce radial motion between rotor and stator This motion varies the air gap flux

density producing stator currents at predictable frequencies fbng that represent the

bearing faults [10]

bng e vf f m f= plusmn sdot (2)

where fe is the electrical supply frequency m=123hellip and fv is one of the characteristics

frequencies listed in the equations of (1)The bearing damage must be significant to

produce noticeable current fluctuations and it may be difficult to distinguish from current

fluctuations due to voltage imbalances machine saturation or statorarmature faults [2]

7

8

212 TEMPERATURE SENSORS

Temperature sensors are largely used in high temperature high load and high speed

environments typically turbine engines and industrial tools Thermocouple sensors

consist of two dissimilar metals joined together at two junctions and when exposed to a

temperature difference generate a thermoelectric voltage dependant only on the material

properties and junction temperature difference [11] In turbine engines these sensors can

be placed on bearing casings or in the lubricant to monitor temperature fluctuations due

to increased friction The thermocouples are fabricated using bulk materials or micro-

machined thin films (TFTC) with the latter providing a faster response small package

and substantially lower production costs [12] The TFTC can be embedded for in situ

monitoring of processes in hostile environments One study successfully embedded

TFTC into electroplated nickel and attained device characteristics on par with surface

mounted devices [13] More recently MEMS temperature sensors have been developed

that consist of micro-machined semiconductor material that undergoes structural

deformation with temperature changes A common implementation is a bent beam

structure whose temperature induced deflection δ is expressed as [14]

( ) 3

2 2

sin12 cos sinz

EA T LEI AEL

α βδ 2β β

∆=

+

(3)

Temperature sensors are well suited for localized fault monitoring and detection but can

only characterize a mechanical system consisting of multiple components through the use

of multiple sensors

9

213 ACOUSTIC EMISSION SENSORS

Acoustic emission (AE) sensors have been used to characterize wear in machine

tools and monitor bearing and gear problems in centrifugal pumps [1516] First

developed as a Non-Destructive Testing (NDT) technique to detect cracks in civil

structures these sensors detect acoustic emissions generated by the release of vibration

waves in a crystalline lattice due to plastic deformation or crack growth [16]

Measurements are made using piezoelectric transducers with high natural frequencies

100 kHz to 1 MHz to capture the ultrasonic AE emissions An AE sensor is useful as it

has the ability to detect subsurface cracks in gear teeth or bearings before appearing on

the surface causing further damage [17] More recently MEMS acoustic sensors have

been developed and one design includes multiple transducers on a single substrate

which each detect acoustic emission energy at different frequencies [18] This helps

distinguish spurious acoustic emissions arising from impact and friction from those

arising from plastic deformation When compared to typical piezoelectric sensors the

MEMS devices have lower sensitivities and fail to detect some acoustic emissions

[1819] In addition the acoustic emission signal suffers severe attenuation as is crosses

various interfaces such as a gearbox or bearing casings In one experiment consisting of

a pinion gear and an associated bearing a 44dB attenuation was seen between an AE

sensor placed directly on the pinion to one placed on the bearing casing and in some

cases intermediate loss of the signal was observed [20]

10

22 MEMS VIBRATION SENSORS

This thesis focuses on the development of a MEMS vibration sensor for fault

detection as the vibration signature of a mechanical system has the potential to give an

overview of the entire system Many types of MEMS vibration sensors exist with

different operating principals based on resonance the piezoelectric effect and

displacement variation

221 RESONANT SENSORS

Resonant sensors are based on the resonant frequency shift of a fixed-fixed micro-

beam that is excited into resonance using electrostatic thermal or piezoelectric methods

Under an applied axial load there is a shift in the resonant frequency which is expressed

as [21]

2 2

2

1 12 1

nn n

EI Fl2

fl A Eα γ

π ρ= +

I

(4)

where l is the beam length E is Youngrsquos Modulus I is the moment of inertia ρ is the

density A is the cross section area F is the applied axial force αn and γn are mode

dependant coefficients and are 4730 and 0295 for mode 1 in the case discussed The

axial force F is generated by a proof mass attached to one beam end which moves in

response to an external acceleration The performance of resonant sensors are dependant

on the quality factor Q which is the ratio of total energy stored in the system to the

energy lost per cycle and determines the sharpness and amplitude of the resonance peak

A high Q resonator develops a peak that is easily distinguishable in a phase locked loop

control circuit and allows for improved performance and resolution [3] A resonator Q

value is lowered by energy losses to the surrounding fluid attached supports and

11

material Losses to the fluid surrounding the MEMS structure is the dominant loss

mechanism for MEMS sensors operated in air and can be greatly reduced by vacuum

packaging which can be accomplished using glass-silicon anodic bonding [22] In

addition to increasing fabrication cost the long term stability of the vacuum degrades

with ultimate device failure stemming from leakage through micro cracks and defects

and out-gassing [22] Support losses arise from the restoring forces they generate and

can be reduce by balanced designs such as a double ended tuning fork (DETF) or

operation in higher modes [3] In addition the sensitivity of a resonant sensor can be

increased by the addition of micro-levers to increase the axial force In one such design a

two stage micro-lever increased the sensitivity by an order of magnitude over a more

conventional single lever arm [2324]

222 PIEZOELECTRIC SENSORS

Piezoelectric materials produce a charge in response to an applied force This is an

intrinsic effect in materials such as Zinc Oxide (ZnO) whereas materials such as Barium

Titanate or Lead Zirconate Titanate (PZT) need to be poled by placing the material in a

strong electric field at an elevated temperature [3] The piezoelectric effect is expressed

mathematical by [25]

i ij j ik kD d T Eε= + (5)

where D represents the electrical displacement (Cm2) d is the charge coefficient (CN)

T is the applied stress matrix (Nm2) ε is the permittivity matrix (Fm) and E is the

electric field matrix (Vm) The subscripts i=123 j=123456 and k=123 represent

the direction to which the physical property is related For micro-sensors poled in the out

of plane direction only the D3 term is relevant and if fabricated using ZnO or PZT only

12

the charge coefficients d31 d32 and d33 are relevant [25] The piezoelectric material is

commonly deposited as a thin film on the surface of compliant structures in bulk micro-

machined sensors or incorporated as a layer in surface micro-machining on a similar

structure [26] However the inclusion of multiple materials introduces large temperature

variations causing out of plane bending especially in surface micro-machined devices

[2627] Poled piezoelectric materials could be depolarized which serves to reduce the

piezoelectric affect in the material This can occur if the sensor is exposed to a strong

electric field of the opposite polarity high temperatures in excess of the Curie point or

high mechanical stress Also the piezoelectric effect reduces over time an effect that can

be reduced with the addition of composite elements [28] Foundry processes such as

PolyMUMPS and SoiMUMPS do not incorporate piezoelectric layers thus sensors using

this method of transduction must use a customized process The piezoelectric sensor

outlined in [26] uses polysilicon surface micro-machining similar to PolyMUMPS

however it incorporates a ZnO layer that is not available in the foundry process

223 DISPLACEMENT VARIATION SENSORS

Displacement variation sensors consist of a proof mass connected to a compliant

micro-flexure The movement of the proof mass is sensed using optical electron

tunneling or more commonly capacitive methods A recent optical accelerometer

utilized a novel multilayer nano-grating for in-plane displacement sensing of the proof

mass [29] For small space variations between the nano-gratings a large change in the

optical reflection amplitude of the grating was observed [30] Even though optical

sensors offer resolutions approaching the Brownian noise limit they are not available in

small packages that are easily placed in space constricted areas for a low cost [29]

13

The electron tunneling effect is observed when the proximity between two electrodes

one flexible and the other fixed is on the order of 10Ǻ [31] If the two electrodes are

biased with voltage VB a tunneling current It flows between the gap [31]

exp( )t B I tgI V xαinfin sdot minus Φ (6)

where α1 is a constant and equal to 1025 Ǻ-1 eV-frac12 Ф is the effective height of the

tunneling barrier and xtg is the tunneling gap When the flexible electrode moves in

response to acceleration a feedback controller maintains the original distance while

determining the applied acceleration Electron tunneling sensors are able to sense

accelerations in the microg-ng range due to the exponential relation between tunneling current

and gap However these highly sensitive devices are susceptible to thermo-mechanical

noise which could be reduced by operation at low pressure [32] It is also important for

the two electrodes to be coated with metal which is difficult to achieve with foundry

fabrication

Capacitive sensors are based on the parallel plate capacitor and are implemented

using inter-digitated comb fingers in an in-plane lateral or transverse drive or in an out of

plane parallel plate drive [33] These sensors have the advantage of high sensitivity low

power consumption small package low temperature dependence and are easily

integrated with existing foundry fabrication processes [34] The lateral drive offers a fully

linear response but low sensitivity whereas the transverse and parallel plate drives

achieve high sensitivity if gap distances are of a few microns and large capacitance areas

are utilized However both drives have a non-linear response and are prone to pull-in

Also if these sensors are fabricated using a surface micromachining process such as

PolyMUMPS where the ground plane is a few microns above the structural layer they

14

are prone to electrical failure due to shorting This was observed in an experiment where

similar sensors were placed in a vibrating environment with a peak of 120g and shorting

between the structural layer and ground plane was found to be the predominate mode of

failure [35] However with certain fabrication processes such as SoiMUMPS this issue

can be neglected as the substrate material below any moving component is removed

Most capacitance sensors are fabricated using an integrated CMOS (Complementary

Metal Oxide Semiconductor)-MEMS technique where the MEMS sensor is fabricated

before during or after the CMOS circuit fabrication [36] Although these sensors offer

high-sensitivity and small parasitic capacitance they suffer from in-plane and out of

plane curling of the beam which reduces the capacitance between adjacent comb fingers

[37] In addition multiple processing steps are required for device fabrication

More than the other transduction methods MEMS capacitive displacement sensors are

attractive for practical implementation in an automobile and therefore is the selected

method for the vibration sensor presented in this work Their low temperature dependence

is an ideal characteristic for an automobile environment that faces varying temperature If

implemented with off-chip readout circuitry they require a standard foundry fabrication

process which provides a strong platform for mass production and low unit cost

3 SENSOR STRUCTURE

This chapter presents the physical structure of the three MEMS vibration sensors

which includes the displacement sensing capacitive comb drive and the micro-

flexureproof mass system First the capacitive comb drive and the three variations

implemented the transverse drive lateral drive and differential tri-plate drive are

discussed Next the mechanical structure is presented and includes the micro-flexure

selection quality factor and equations of motion Finally instability that places a limit

on the operating range of the sensor as a result of a gap closing capacitive comb drive is

discussed

31 SENSING

Capacitive sensors are based on the parallel plate capacitor and are commonly

implemented using inter-digitated comb fingers A capacitance can be realized if a voltage

is placed between two closely spaced plates If the fringing electric field is neglected the

capacitance of the parallel plate capacitor Cpp (F) is expressed as [28]

air pppp

k AC

dε

= (7)

where kair=1 is the dielectric constant of air at atmospheric pressure ε=88542times10-12 F m-1

is the permittivity of free space App is the overlapping area (m2) and d is the distance

between the two plates (m)

15

16

Figure 3-1 An inter-digitated comb drive where changes in capacitance are generated by either changes in gap distance x0 or in the overlap area y0timest

Figure 3-1

For a set of inter-digitated comb fingers or comb drive shown in the total

capacitance Ctotal is expressed as [28]

0 0

0 0

( ) ( )( ) (Total cy y y yC N t

)x x x xε

+ += + + minus

(8)

where N is the number of comb fingers on the movable comb y0 is the initial comb finger

overlap length (m) tc is the out of plane comb finger thickness (m) and x0 is the inital

finger gap (m) In this work the comb drive is implemented in three arrangements

transverse lateral and differential tri-plate In a transverse comb drive changes in

capacitance are generated by transverse movements along the x-axis resulting in changes

to the gap x0 between the movable and fixed comb fingers The nonlinear relation

between total capacitance CTotalT and displacement is expressed as

17

00 0

1 1( ) (Total TC N ty

)x x x xε

= + minus +

(9)

Although the drive sensitivity part is not constant an average sensitivity can be

assumed for small displacements [38]

Total TC xpart

In a lateral comb drive changes in capacitance are generated by lateral movements

along the y-axis resulting in changes to the overlapping area y0timestc The total capacitance

CTotalL between all comb fingers is expressed as

00

2 ( )Total LN tC yx

yε= +

(10)

The sensitivity of the lateral comb drive SL is expressed as

0

2Total LL

C N tSy x

εpart= =

part

(11)

The lateral comb drive was first introduced by W C Tang T H Nguyen and R T

Howe [33] in an electrostatic resonator and has since been used numerous times in both

drive and sense configurations It has the advantage of a fully linear response and

negligible instability but offers low sensitivity The sensitivity for the lateral drive L1

implemented in this work is 0038 pF microm-1

A differential tri-plate drive shown in Figure 3-2 which is a variation of the

transverse drive is implemented to improve sensitivity and reduce noise [39]

18

Figure 3-2 An inter-digitated comb drive in a differential tri-plate arrangement where x01 x02 and x03 are the gap distances and x01= x02ltlt x03 to improve sensitivity

In the tri-plate drive the at rest gap distances between the two capacitors formed by any

movable finger is not equivalent For a differential arrangement insuring that x01=

x02ltltx03 the linearity for small displacements (less than 1 microm) is enhanced over the case

of a single ended arrangement where x01=x02=x03=x0 which is equivalent to the

transverse drive This is shown graphically in Figure 3-3 where x01= x02=x03=2 microm in

the lower curve and x01= x02= 2 microm and x03= 15 microm in the upper curve A substantial

improvement in linearity is attained with the tri-plate drive

19

0

005

01

015

02

025

03

035

0 02 04 06 08 1

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-3 Comparison of the effect of gap distances on linearity for two case lsquorsquo represents case 1 where the fingers in are implemented in a differential tri-plate arrangement and x01= x02= 2 microm and x03= 15 microm lsquorsquo represents the case 2 where the fingers are implemented in a single ended arrangement and x01= x02=x03=2 microm

The tri-plate drive is also implemented in a differential arrangement where the output

is the change in capacitance rather than the total capacitance resulting from an external

excitation This reduces errors cause by off-axis excitation and common mode noise The

output of the differential tri-plate drive is expressed as [39]

Total Top Total BottomC C C∆ = minus where

003 01

1 1( ) (Total TopC N ty

)x x x xε

= +

minus +

(12)

20

002 03

1 1( ) (Total BottomC N ty

)x x x xε

= +

minus +

A graphic comparison of the capacitancedisplacement relation for sensors T1 and T2 is

shown in Figure 3-4 For these gap closing drives it is assumed that for displacements

less that 1 microm the relation is linear indicated by the dashed line above each curve The

sensitivity within this region is approximately 0046 pF microm-1 and 0253 pF microm-1 for T1

and T2 respectively The order of magnitude improvement is attributed to the differential

tri-plate arrangement and thicker comb finger

0

02

04

06

08

1

12

14

16

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-4 lsquorsquo represents the output of the differential tri-plate drive while lsquorsquo represents the output of the transverse drive The dashed line represents the linear approximation for displacements less that 1 microm For the purpose of comparison the output of the transverse drive was equalized to zero at zero displacement

21

In this work the transverse drive is first implemented in sensor T1 using the

PolyMUMPS fabrication process Then in an effort to linearize the sensor response the

lateral drive is used in L1 using SoiMUMPS The reduction in drive sensitivity due to the

lateral drive is further reduced by increasing the gap distance to 3 microm which is

recommended by the foundry to ensure a successful device These reductions are

balanced by increasing the number of comb fingers N and selecting a substrate thickness

that is 125 times large than T1 The third device T2 is fabricated using SoiMUMPS and

uses the tri-plate differential drive This design capitalizes on the increase in sensitivity of

a gap closing drive and larger substrate thickness attaining an order of magnitude

increase in sensitivity over T1 and L1 The specifications for each drive along with the

successive improvement in drive sensitivity are outlined in Table 3-1

Table 3-1 Specifications for each capacitive comb drive implemented in this work T1 L1 and T2

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Comb Finger Thickness tc (microm) 2 25 25 Gap Distance x0 (microm) 2 3 2 Number of Fingers N 200 256 30 Overlap length y0 (microm) 45 10 70 Drive Sensitivity (pF microm-1) 0046 0038 0273

32 MECHANICAL ANALYSIS

321 MICRO-FLEXURE SELECTION

The sensor structure consists of a proof mass connected to a compliant micro-flexure

that itself is anchored to the sensor substrate The proof mass displaces in response to an

22

applied acceleration and this displacement is sensed using the capacitive comb drive

attached to the proof mass

Proof mass

Folded beam micro-flexure

Anchor

Figure 3-5 Sensor Structure

This section focuses on micro-flexure selection as it should exhibit a large linear

displacement range and have a high lateral stiffness ratio to reduce cross-axis sensitivity

A number of micro-flexures have been used in MEMS research these include the

clamped-clamped crab-leg and folded beam shown in Figure 3-6-through 3-7

The clamped-clamped flexure exhibits a high stiffness ratio however for large lateral

displacements extensional axial forces buildup within the beams resulting in a non-linear

force-displacement relation [40] Within the linear regime the stiffness of the clamped-

clamped flexure is expressed as [40]

3

3b

yEt wk

L=

(13)

23

where E is the Youngrsquos Modulus of the beam material (N m-2) tb is the beam out of

plane thickness (m) w is the beam width (m) and L is the beam length (m) The stiffness

ratio is expressed as [40]

2x

y

k Lk w

=

(14)

If L is two orders of magnitude greater than w this results in a ratio of approximately

10000 The linear displacement range is determined using Finite Element Analysis (FEA)

of a clamped-clamped flexure with L=387 microm w=4 microm and tb=2 microm The result of this

simulation is plotted in and compared to a structure with constant stiffness ky

= 035 N m-1 predicted by Eq (13) The linear relation is valid for small deflections that

are less than 2 microm In addition this micro-flexure is susceptible to buckling as a result of

a compressive residual stress that may arise due to the fabrication process [41]

Figure 3-9

In the crab leg design the thigh section is meant to relieve the extensional axial

stress postponing the occurrence of non-linearity until greater lateral displacement [42]

The stiffness of the crab leg flexure is expressed as [43]

3

44

b s s ty

s s t

Et w L LkL L L

κκ

+= +

(15)

where κ=IsIt=(wswt)3 is a ratio of the second moment of inertia of the shin (s) and thigh

(t) ws and wt represent the width and Ls and Lt represent the length of the shin and thigh

regions respectively The stiffness ratio is expressed as [43]

34

4t s s tx

y s t s

w L L Lkk w L L L

αα t

+= +

(16)

The linear displacement range of a crab leg flexure with Ls=285 microm Lt=30 microm ws=wt=3

microm and tb=2 microm is determined using FEA The results are plotted in Figure 3-9 and

compared to a structure of constant stiffness ky=035 N m-1 predicted by Eq (15) The

24

non-linearity is greatly reduced when compared to the clamp-clamped flexure with Eq

accurately predicting the deflection until 4 microm However the flexure is more

susceptible to off-axis deflection as there is an order of magnitude reduction in the

stiffness ratio as compared to the clamped-clamped flexure

(15)

The folded beam flexure stands out as it is designed to exhibit a large linear

displacement range and to relieve residual stress due to the manufacturing process by

allowing the beams to expand and contract along the axial direction [40] It was first used

in an electrostatic resonator and has subsequently been used for electrostatic actuation in

a magnetic hard drives and micro-positioning [3344] If the truss section is rigid

ItrussgtgtIbeam then the stiffness of the flexure is expressed as [4043]

3

32b

yEt wk

L=

(17)

The stiffness ratio is expressed as [40]

2x

y

k Lk w

=

(18)

Figure 3-6 Clamped-clamped flexure

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

Ls shin beam length Lt thigh beam length m effective proof mass N number of comb fingers on the movable comb

NB number of balls

Qtotal total quality factor

Qf energy lost to surrounding fluid

Qs energy lost through supports

Qm energy lost through sensor material S theoretical sensitivity of sensor SL sensitivity of a lateral comb drive SM theoretical mechanical sensitivity of the proof massflexure system SRO sensitivity of readout circuit ST sensitivity of a transverse comb drive STAVG average sensitivity of a transverse comb drive t time T absolute temperature tb flexure beam thickness tc out of plane comb finger thickness

Vact electrostatic actuation voltage VCA output amplitude of charge amplifier VDM output amplitide of demodulator VE input amplitude of sinusoidal excitation voltage VLP output of low pass filter VLPMAX maximum output voltage at low pass filter VVA output amplitude of voltage amplifier vy relative motion of proof mass with respect to base Vy relative motion amplitude w flexure beam width Wplate effective plate width ws shin beam width wt thigh beam width x0 initial comb finger gap

x01 small gap distance on top side of the tri-plate comb drive

x02 small gap distance on bottom side of the tri-plate comb drive

ix

x03 larger gap distance on the tri-plate comb drive

XN noise displacement y0 initial comb finger overlap length yB base displacement YB base displacement amplitude ym absolute proof mass displacement α parameter dependant on the WplateLplate ratio ∆C Change in capacitance from a tri-plate drive in a differential

arrangement ε permittivity of free space κ second moment of inertia ratio (shinthigh) ζ damping ratio microair absolute viscosity of air ω base excitation frequency in rads ωn sensor natural frequency in rads ωv electrostatic actuation voltage frequency in rads L1 Sensor fabricated using SoiMUMPS with a lateral comb drive T1 Sensor fabricated using PolyMUMPS with a transverse comb drive T2 Sensor fabricated using SoiMUMPS with a tri-plate comb drive

x

1 INTRODUCTION

11 RESEARCH PURPOSE

The modern automobile consists of many mechanical systems such as power seats

windshield wipers mirrors trunks and windows which are all susceptible to breakdown

Without any condition monitoring system the breakdown is usually catastrophic and

requires an expensive part replacement Real-time condition monitoring allows for early

detection of faults which could require a simple solution such as the application of a

lubricant to fix This prolongs the useful life of the component and prevents sudden and

unexpected failure Real-time condition monitoring can be accomplished by examining the

vibration signature of a mechanical system For example an automobile power window

consists of a DC motor and its associated bearings and couplings a gear reduction system

consisting of worm and spur gears and kinematic links Faults resulting in excessive

vibrations may be caused by coupling misalignment bearing failure or gear train failure

Coupling misalignments occur at the connection between the drive shaft and the driven

shaft and are typically due to imperfect manufacturing Bearing failure is caused by lack

of lubrication or moisture contamination causing rusting while gear train failure is caused

by misaligned broken cracked or chipped gear teeth [1] Each fault occurs at a

characteristic frequency and so the state of the mechanical system can be determined by

monitoring the amplitudes of the relevant frequencies Vibrations due to coupling

misalignments occur at harmonics of the shaft rotational speed Gear vibrations occur at

the gear turn speed or at sidebands of the gear mesh frequency [1] Ball bearing vibrations

1

2

may be caused by outer bearing race defects fOB inner bearing race defects fIB ball

defects fB and train defects fT which all occur at specific frequencies [12]

cos( )12

dBOB r

p

bNf fd

β = minus

cos( )12

dBIB r

p

bNf fd

β = +

2cos( )1

2p r d

Bd p

d f bfb d

β = minus

cos( )12

drT

p

bffd

β = minus

(1)

where NB is the number of balls fr is the rotation frequency bd is the ball diameter β is the

contact angle between the ball and races and dp is the ball pitch diameter By monitoring

the real time vibration signature of the mechanical system anomalies can be quickly

identified and fixed

12 MICROELECTROMECHANICAL SENSORS

Microelectromechanical systems (MEMS) inertial sensors provide a small footprint

with sensitivities that are either comparable or exceed any macro sensor along with the

capability of mass production and low unit cost These sensors utilize compliant micro-

flexures attached to a proof mass that displaces in response to an environmental

acceleration Many transduction mechanisms have been developed that convert the

displacement into a measurable electric signal and include thermal piezoresistive

piezoelectric optical and capacitive methods Most MEMS sensors are silicon based and

are fabricated using surface or bulk micromachining Surface micromachining creates free

standing movable structures on top of a substrate using a combination of sacrificial layers

3

and structural layers which are commonly polysilicon [3] In bulk micromachining the

mechanical structures are defined using a removal process where bulk material typically

silicon is etched away MEMS sensors have been used for vibration and shock monitoring

on industrial systems and robotics guidance and navigation in global positioning systems

(GPS) seismometry in earthquake prediction image stabilization in digital cameras and

automobile safety and stability [4]

13 AUTOMOTIVE SENSORS

Sensors cover every major aspect of a modern automobile power-train sensors

monitor fuel combustion and emissions chassis sensors monitor road traction and tire

condition and body sensors facilitate air-bag deployment and vehicle proximity for radar

guided cruise control [5] Pressure sensors which typically consist of a piezoresistive

strain sensing element attached to a silicon diaphragm that deflects when exposed to an

applied pressure are one of the first micro-machined sensors used in an automobile

Implemented as manifold absolute pressure (MAP) sensors they allow precise control of

the air fuel ratio which allows the catalytic converter to efficiently reduce tailpipe

emissions [6] Variable reluctance sensors based on electromagnetics are used for

automobile traction control and produce a voltage output that is dependant on the

magnetic flux variations between a rotating component and the sensor bias magnet [57]

MEMS based linear accelerometers are utilized for airbag deployment upon impact and

provide a lower cost and smaller package over the older ball-in-tube accelerometers [6]

4

14 OPERATING ENVIRONMENT

Through proper packaging MEMS sensors can withstand the harsh automotive

environment where they face cyclic temperatures (-40degC to 150degC) mechanical shock or

vibration (10g-500g) and exposure to various fluids (oil brake fluid or engine coolant)

and airborne particles [8] Temperature changes induce thermal stresses causing

undesirable bending of the silicon microstructures that changes their dynamic response

Mechanical shock which could occur during a high impact collision or vibration from

uneven road conditions cause the delicate microstructures to fracture at points of stress

concentration These effects could be reduced by efficient design Use of a folded beam

micro-flexure that allows expansion and contraction or use of fillets at corners which

reduce the chance of fracture Bonding pads wire bonding and onboard circuit

components can be protected from air borne particles which could cause shorting using

parylene or gels Alternatively the entire package can be sealed hermetically which also

protects the MEMS device [8] In some cases such as fluid pressure sensors intimate

contact between the MEMS device and potentially corrosive fluid is necessary A recent

approach used a nanometer thin coating of silicon carbide (SiC) over a foundry fabricated

MEMS sensor using low temperature low pressure chemical vapor deposition (LPCVD)

[9] The SiC known for its chemical inertness did not compromise the mechanical

characteristics of the original design Therefore MEMS provides a strong platform for an

automotive sensor

5

15 OBJECTIVES AND CONTRIBUTION

The objective of this work is the development of a low frequency MEMS vibration

sensor The sensor must be low cost and mass producible which can be achieved using

standard foundry fabrication processes such as the Polysilicon Multi-User MEMS

Processes (PolyMUMPS) or the Silicon-on-Insulator Multi-User MEMS Processes

(SoiMUMPS) In addition the number of additional processing steps which are required

if integrated CMOS circuitry is used or if vacuum packaging is required must be

minimized This can be achieved using an off-chip circuit made with widely available

discrete components and sensor operation in air

In addition this work presents a MEMS sensor that is fabricated using standard

foundry processes such as PolyMUMPS and SoiMUMPS compared to many MEMS

sensors in literature that are fabricated using customized processes

16 RESEARCH OUTLINE

This work demonstrates three iterations of a low frequency capacitive vibration

sensor fabricated using the standard foundry processes Each sensor consists of a

suspended proof mass that displaces in response to an external vibration The proof mass

is attached to a folded beam micro-flexure compliant in one direction and its lateral

movement is sensed using a capacitive comb drive The first sensor T1 is fabricated in

PolyMUMPS and uses a transverse capacitive comb drive The second L1 fabricated in

SoiMUMPS uses a fully linear lateral capacitive drive and achieves an increase in

sensitivity from T1 Finally the third sensor T2 is fabricated in SoiMUMPS and uses a

differential tri-plate capacitive drive This sensor takes advantage of the increase in

6

sensitivity gained from a transverse drive but rejects off-axis displacement errors and

common mode using a differential tri-plate arrangement In addition an off-chip

capacitance to voltage readout circuit is fabricated and tested on sensor L1

Chapter 2 begins with an overview on the various methods of fault detection

followed by a review of MEMS vibration sensors each with a different transduction

mechanism Chapter 3 outlines the sensor structure focusing on the mechanical design

and sensing principal Chapter 4 focuses on the readout circuitry layout and simulation

results Chapter 5 outlines the various source of noise including mechanical noise arising

from the sensor and electric noise from the readout circuit Chapter 6 overviews the

foundry fabrication processes used including PolyMUMPs and SoiMUMPS The

experimental results of two sensors T1 and L1 which were fabricated and tested within

the thesis time frame are presented in Chapter 7 Chapter 8 concludes the thesis with a

brief summary and outline for future work

2 LITERATURE REVIEW

21 EXISTING METHODS FOR FAULT DETECTION

As this thesis develops a sensor for fault detection using vibration monitoring this

section overviews three existing methods for fault detection in mechanical systems First

motor current monitoring is discussed as it has the ability to measure significant bearing

faults in motor systems Next temperature sensors specifically thin film thermocouples

(TFTC) and MEMS temperature sensors are discussed These sensors are mounted in the

chamber of a lubricating fluid or in close proximity to a moving component and monitor

temperature fluctuations caused by increased wear Finally acoustic emission (AE)

sensors are discussed which have the ability to detect ultrasonic emissions generated by

plastic deformation in the crystalline lattice of a mechanical component

211 MOTOR CURRENT MONITORING

Monitoring of stator current is used as a method of determining bearing related faults

that produce radial motion between rotor and stator This motion varies the air gap flux

density producing stator currents at predictable frequencies fbng that represent the

bearing faults [10]

bng e vf f m f= plusmn sdot (2)

where fe is the electrical supply frequency m=123hellip and fv is one of the characteristics

frequencies listed in the equations of (1)The bearing damage must be significant to

produce noticeable current fluctuations and it may be difficult to distinguish from current

fluctuations due to voltage imbalances machine saturation or statorarmature faults [2]

7

8

212 TEMPERATURE SENSORS

Temperature sensors are largely used in high temperature high load and high speed

environments typically turbine engines and industrial tools Thermocouple sensors

consist of two dissimilar metals joined together at two junctions and when exposed to a

temperature difference generate a thermoelectric voltage dependant only on the material

properties and junction temperature difference [11] In turbine engines these sensors can

be placed on bearing casings or in the lubricant to monitor temperature fluctuations due

to increased friction The thermocouples are fabricated using bulk materials or micro-

machined thin films (TFTC) with the latter providing a faster response small package

and substantially lower production costs [12] The TFTC can be embedded for in situ

monitoring of processes in hostile environments One study successfully embedded

TFTC into electroplated nickel and attained device characteristics on par with surface

mounted devices [13] More recently MEMS temperature sensors have been developed

that consist of micro-machined semiconductor material that undergoes structural

deformation with temperature changes A common implementation is a bent beam

structure whose temperature induced deflection δ is expressed as [14]

( ) 3

2 2

sin12 cos sinz

EA T LEI AEL

α βδ 2β β

∆=

+

(3)

Temperature sensors are well suited for localized fault monitoring and detection but can

only characterize a mechanical system consisting of multiple components through the use

of multiple sensors

9

213 ACOUSTIC EMISSION SENSORS

Acoustic emission (AE) sensors have been used to characterize wear in machine

tools and monitor bearing and gear problems in centrifugal pumps [1516] First

developed as a Non-Destructive Testing (NDT) technique to detect cracks in civil

structures these sensors detect acoustic emissions generated by the release of vibration

waves in a crystalline lattice due to plastic deformation or crack growth [16]

Measurements are made using piezoelectric transducers with high natural frequencies

100 kHz to 1 MHz to capture the ultrasonic AE emissions An AE sensor is useful as it

has the ability to detect subsurface cracks in gear teeth or bearings before appearing on

the surface causing further damage [17] More recently MEMS acoustic sensors have

been developed and one design includes multiple transducers on a single substrate

which each detect acoustic emission energy at different frequencies [18] This helps

distinguish spurious acoustic emissions arising from impact and friction from those

arising from plastic deformation When compared to typical piezoelectric sensors the

MEMS devices have lower sensitivities and fail to detect some acoustic emissions

[1819] In addition the acoustic emission signal suffers severe attenuation as is crosses

various interfaces such as a gearbox or bearing casings In one experiment consisting of

a pinion gear and an associated bearing a 44dB attenuation was seen between an AE

sensor placed directly on the pinion to one placed on the bearing casing and in some

cases intermediate loss of the signal was observed [20]

10

22 MEMS VIBRATION SENSORS

This thesis focuses on the development of a MEMS vibration sensor for fault

detection as the vibration signature of a mechanical system has the potential to give an

overview of the entire system Many types of MEMS vibration sensors exist with

different operating principals based on resonance the piezoelectric effect and

displacement variation

221 RESONANT SENSORS

Resonant sensors are based on the resonant frequency shift of a fixed-fixed micro-

beam that is excited into resonance using electrostatic thermal or piezoelectric methods

Under an applied axial load there is a shift in the resonant frequency which is expressed

as [21]

2 2

2

1 12 1

nn n

EI Fl2

fl A Eα γ

π ρ= +

I

(4)

where l is the beam length E is Youngrsquos Modulus I is the moment of inertia ρ is the

density A is the cross section area F is the applied axial force αn and γn are mode

dependant coefficients and are 4730 and 0295 for mode 1 in the case discussed The

axial force F is generated by a proof mass attached to one beam end which moves in

response to an external acceleration The performance of resonant sensors are dependant

on the quality factor Q which is the ratio of total energy stored in the system to the

energy lost per cycle and determines the sharpness and amplitude of the resonance peak

A high Q resonator develops a peak that is easily distinguishable in a phase locked loop

control circuit and allows for improved performance and resolution [3] A resonator Q

value is lowered by energy losses to the surrounding fluid attached supports and

11

material Losses to the fluid surrounding the MEMS structure is the dominant loss

mechanism for MEMS sensors operated in air and can be greatly reduced by vacuum

packaging which can be accomplished using glass-silicon anodic bonding [22] In

addition to increasing fabrication cost the long term stability of the vacuum degrades

with ultimate device failure stemming from leakage through micro cracks and defects

and out-gassing [22] Support losses arise from the restoring forces they generate and

can be reduce by balanced designs such as a double ended tuning fork (DETF) or

operation in higher modes [3] In addition the sensitivity of a resonant sensor can be

increased by the addition of micro-levers to increase the axial force In one such design a

two stage micro-lever increased the sensitivity by an order of magnitude over a more

conventional single lever arm [2324]

222 PIEZOELECTRIC SENSORS

Piezoelectric materials produce a charge in response to an applied force This is an

intrinsic effect in materials such as Zinc Oxide (ZnO) whereas materials such as Barium

Titanate or Lead Zirconate Titanate (PZT) need to be poled by placing the material in a

strong electric field at an elevated temperature [3] The piezoelectric effect is expressed

mathematical by [25]

i ij j ik kD d T Eε= + (5)

where D represents the electrical displacement (Cm2) d is the charge coefficient (CN)

T is the applied stress matrix (Nm2) ε is the permittivity matrix (Fm) and E is the

electric field matrix (Vm) The subscripts i=123 j=123456 and k=123 represent

the direction to which the physical property is related For micro-sensors poled in the out

of plane direction only the D3 term is relevant and if fabricated using ZnO or PZT only

12

the charge coefficients d31 d32 and d33 are relevant [25] The piezoelectric material is

commonly deposited as a thin film on the surface of compliant structures in bulk micro-

machined sensors or incorporated as a layer in surface micro-machining on a similar

structure [26] However the inclusion of multiple materials introduces large temperature

variations causing out of plane bending especially in surface micro-machined devices

[2627] Poled piezoelectric materials could be depolarized which serves to reduce the

piezoelectric affect in the material This can occur if the sensor is exposed to a strong

electric field of the opposite polarity high temperatures in excess of the Curie point or

high mechanical stress Also the piezoelectric effect reduces over time an effect that can

be reduced with the addition of composite elements [28] Foundry processes such as

PolyMUMPS and SoiMUMPS do not incorporate piezoelectric layers thus sensors using

this method of transduction must use a customized process The piezoelectric sensor

outlined in [26] uses polysilicon surface micro-machining similar to PolyMUMPS

however it incorporates a ZnO layer that is not available in the foundry process

223 DISPLACEMENT VARIATION SENSORS

Displacement variation sensors consist of a proof mass connected to a compliant

micro-flexure The movement of the proof mass is sensed using optical electron

tunneling or more commonly capacitive methods A recent optical accelerometer

utilized a novel multilayer nano-grating for in-plane displacement sensing of the proof

mass [29] For small space variations between the nano-gratings a large change in the

optical reflection amplitude of the grating was observed [30] Even though optical

sensors offer resolutions approaching the Brownian noise limit they are not available in

small packages that are easily placed in space constricted areas for a low cost [29]

13

The electron tunneling effect is observed when the proximity between two electrodes

one flexible and the other fixed is on the order of 10Ǻ [31] If the two electrodes are

biased with voltage VB a tunneling current It flows between the gap [31]

exp( )t B I tgI V xαinfin sdot minus Φ (6)

where α1 is a constant and equal to 1025 Ǻ-1 eV-frac12 Ф is the effective height of the

tunneling barrier and xtg is the tunneling gap When the flexible electrode moves in

response to acceleration a feedback controller maintains the original distance while

determining the applied acceleration Electron tunneling sensors are able to sense

accelerations in the microg-ng range due to the exponential relation between tunneling current

and gap However these highly sensitive devices are susceptible to thermo-mechanical

noise which could be reduced by operation at low pressure [32] It is also important for

the two electrodes to be coated with metal which is difficult to achieve with foundry

fabrication

Capacitive sensors are based on the parallel plate capacitor and are implemented

using inter-digitated comb fingers in an in-plane lateral or transverse drive or in an out of

plane parallel plate drive [33] These sensors have the advantage of high sensitivity low

power consumption small package low temperature dependence and are easily

integrated with existing foundry fabrication processes [34] The lateral drive offers a fully

linear response but low sensitivity whereas the transverse and parallel plate drives

achieve high sensitivity if gap distances are of a few microns and large capacitance areas

are utilized However both drives have a non-linear response and are prone to pull-in

Also if these sensors are fabricated using a surface micromachining process such as

PolyMUMPS where the ground plane is a few microns above the structural layer they

14

are prone to electrical failure due to shorting This was observed in an experiment where

similar sensors were placed in a vibrating environment with a peak of 120g and shorting

between the structural layer and ground plane was found to be the predominate mode of

failure [35] However with certain fabrication processes such as SoiMUMPS this issue

can be neglected as the substrate material below any moving component is removed

Most capacitance sensors are fabricated using an integrated CMOS (Complementary

Metal Oxide Semiconductor)-MEMS technique where the MEMS sensor is fabricated

before during or after the CMOS circuit fabrication [36] Although these sensors offer

high-sensitivity and small parasitic capacitance they suffer from in-plane and out of

plane curling of the beam which reduces the capacitance between adjacent comb fingers

[37] In addition multiple processing steps are required for device fabrication

More than the other transduction methods MEMS capacitive displacement sensors are

attractive for practical implementation in an automobile and therefore is the selected

method for the vibration sensor presented in this work Their low temperature dependence

is an ideal characteristic for an automobile environment that faces varying temperature If

implemented with off-chip readout circuitry they require a standard foundry fabrication

process which provides a strong platform for mass production and low unit cost

3 SENSOR STRUCTURE

This chapter presents the physical structure of the three MEMS vibration sensors

which includes the displacement sensing capacitive comb drive and the micro-

flexureproof mass system First the capacitive comb drive and the three variations

implemented the transverse drive lateral drive and differential tri-plate drive are

discussed Next the mechanical structure is presented and includes the micro-flexure

selection quality factor and equations of motion Finally instability that places a limit

on the operating range of the sensor as a result of a gap closing capacitive comb drive is

discussed

31 SENSING

Capacitive sensors are based on the parallel plate capacitor and are commonly

implemented using inter-digitated comb fingers A capacitance can be realized if a voltage

is placed between two closely spaced plates If the fringing electric field is neglected the

capacitance of the parallel plate capacitor Cpp (F) is expressed as [28]

air pppp

k AC

dε

= (7)

where kair=1 is the dielectric constant of air at atmospheric pressure ε=88542times10-12 F m-1

is the permittivity of free space App is the overlapping area (m2) and d is the distance

between the two plates (m)

15

16

Figure 3-1 An inter-digitated comb drive where changes in capacitance are generated by either changes in gap distance x0 or in the overlap area y0timest

Figure 3-1

For a set of inter-digitated comb fingers or comb drive shown in the total

capacitance Ctotal is expressed as [28]

0 0

0 0

( ) ( )( ) (Total cy y y yC N t

)x x x xε

+ += + + minus

(8)

where N is the number of comb fingers on the movable comb y0 is the initial comb finger

overlap length (m) tc is the out of plane comb finger thickness (m) and x0 is the inital

finger gap (m) In this work the comb drive is implemented in three arrangements

transverse lateral and differential tri-plate In a transverse comb drive changes in

capacitance are generated by transverse movements along the x-axis resulting in changes

to the gap x0 between the movable and fixed comb fingers The nonlinear relation

between total capacitance CTotalT and displacement is expressed as

17

00 0

1 1( ) (Total TC N ty

)x x x xε

= + minus +

(9)

Although the drive sensitivity part is not constant an average sensitivity can be

assumed for small displacements [38]

Total TC xpart

In a lateral comb drive changes in capacitance are generated by lateral movements

along the y-axis resulting in changes to the overlapping area y0timestc The total capacitance

CTotalL between all comb fingers is expressed as

00

2 ( )Total LN tC yx

yε= +

(10)

The sensitivity of the lateral comb drive SL is expressed as

0

2Total LL

C N tSy x

εpart= =

part

(11)

The lateral comb drive was first introduced by W C Tang T H Nguyen and R T

Howe [33] in an electrostatic resonator and has since been used numerous times in both

drive and sense configurations It has the advantage of a fully linear response and

negligible instability but offers low sensitivity The sensitivity for the lateral drive L1

implemented in this work is 0038 pF microm-1

A differential tri-plate drive shown in Figure 3-2 which is a variation of the

transverse drive is implemented to improve sensitivity and reduce noise [39]

18

Figure 3-2 An inter-digitated comb drive in a differential tri-plate arrangement where x01 x02 and x03 are the gap distances and x01= x02ltlt x03 to improve sensitivity

In the tri-plate drive the at rest gap distances between the two capacitors formed by any

movable finger is not equivalent For a differential arrangement insuring that x01=

x02ltltx03 the linearity for small displacements (less than 1 microm) is enhanced over the case

of a single ended arrangement where x01=x02=x03=x0 which is equivalent to the

transverse drive This is shown graphically in Figure 3-3 where x01= x02=x03=2 microm in

the lower curve and x01= x02= 2 microm and x03= 15 microm in the upper curve A substantial

improvement in linearity is attained with the tri-plate drive

19

0

005

01

015

02

025

03

035

0 02 04 06 08 1

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-3 Comparison of the effect of gap distances on linearity for two case lsquorsquo represents case 1 where the fingers in are implemented in a differential tri-plate arrangement and x01= x02= 2 microm and x03= 15 microm lsquorsquo represents the case 2 where the fingers are implemented in a single ended arrangement and x01= x02=x03=2 microm

The tri-plate drive is also implemented in a differential arrangement where the output

is the change in capacitance rather than the total capacitance resulting from an external

excitation This reduces errors cause by off-axis excitation and common mode noise The

output of the differential tri-plate drive is expressed as [39]

Total Top Total BottomC C C∆ = minus where

003 01

1 1( ) (Total TopC N ty

)x x x xε

= +

minus +

(12)

20

002 03

1 1( ) (Total BottomC N ty

)x x x xε

= +

minus +

A graphic comparison of the capacitancedisplacement relation for sensors T1 and T2 is

shown in Figure 3-4 For these gap closing drives it is assumed that for displacements

less that 1 microm the relation is linear indicated by the dashed line above each curve The

sensitivity within this region is approximately 0046 pF microm-1 and 0253 pF microm-1 for T1

and T2 respectively The order of magnitude improvement is attributed to the differential

tri-plate arrangement and thicker comb finger

0

02

04

06

08

1

12

14

16

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-4 lsquorsquo represents the output of the differential tri-plate drive while lsquorsquo represents the output of the transverse drive The dashed line represents the linear approximation for displacements less that 1 microm For the purpose of comparison the output of the transverse drive was equalized to zero at zero displacement

21

In this work the transverse drive is first implemented in sensor T1 using the

PolyMUMPS fabrication process Then in an effort to linearize the sensor response the

lateral drive is used in L1 using SoiMUMPS The reduction in drive sensitivity due to the

lateral drive is further reduced by increasing the gap distance to 3 microm which is

recommended by the foundry to ensure a successful device These reductions are

balanced by increasing the number of comb fingers N and selecting a substrate thickness

that is 125 times large than T1 The third device T2 is fabricated using SoiMUMPS and

uses the tri-plate differential drive This design capitalizes on the increase in sensitivity of

a gap closing drive and larger substrate thickness attaining an order of magnitude

increase in sensitivity over T1 and L1 The specifications for each drive along with the

successive improvement in drive sensitivity are outlined in Table 3-1

Table 3-1 Specifications for each capacitive comb drive implemented in this work T1 L1 and T2

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Comb Finger Thickness tc (microm) 2 25 25 Gap Distance x0 (microm) 2 3 2 Number of Fingers N 200 256 30 Overlap length y0 (microm) 45 10 70 Drive Sensitivity (pF microm-1) 0046 0038 0273

32 MECHANICAL ANALYSIS

321 MICRO-FLEXURE SELECTION

The sensor structure consists of a proof mass connected to a compliant micro-flexure

that itself is anchored to the sensor substrate The proof mass displaces in response to an

22

applied acceleration and this displacement is sensed using the capacitive comb drive

attached to the proof mass

Proof mass

Folded beam micro-flexure

Anchor

Figure 3-5 Sensor Structure

This section focuses on micro-flexure selection as it should exhibit a large linear

displacement range and have a high lateral stiffness ratio to reduce cross-axis sensitivity

A number of micro-flexures have been used in MEMS research these include the

clamped-clamped crab-leg and folded beam shown in Figure 3-6-through 3-7

The clamped-clamped flexure exhibits a high stiffness ratio however for large lateral

displacements extensional axial forces buildup within the beams resulting in a non-linear

force-displacement relation [40] Within the linear regime the stiffness of the clamped-

clamped flexure is expressed as [40]

3

3b

yEt wk

L=

(13)

23

where E is the Youngrsquos Modulus of the beam material (N m-2) tb is the beam out of

plane thickness (m) w is the beam width (m) and L is the beam length (m) The stiffness

ratio is expressed as [40]

2x

y

k Lk w

=

(14)

If L is two orders of magnitude greater than w this results in a ratio of approximately

10000 The linear displacement range is determined using Finite Element Analysis (FEA)

of a clamped-clamped flexure with L=387 microm w=4 microm and tb=2 microm The result of this

simulation is plotted in and compared to a structure with constant stiffness ky

= 035 N m-1 predicted by Eq (13) The linear relation is valid for small deflections that

are less than 2 microm In addition this micro-flexure is susceptible to buckling as a result of

a compressive residual stress that may arise due to the fabrication process [41]

Figure 3-9

In the crab leg design the thigh section is meant to relieve the extensional axial

stress postponing the occurrence of non-linearity until greater lateral displacement [42]

The stiffness of the crab leg flexure is expressed as [43]

3

44

b s s ty

s s t

Et w L LkL L L

κκ

+= +

(15)

where κ=IsIt=(wswt)3 is a ratio of the second moment of inertia of the shin (s) and thigh

(t) ws and wt represent the width and Ls and Lt represent the length of the shin and thigh

regions respectively The stiffness ratio is expressed as [43]

34

4t s s tx

y s t s

w L L Lkk w L L L

αα t

+= +

(16)

The linear displacement range of a crab leg flexure with Ls=285 microm Lt=30 microm ws=wt=3

microm and tb=2 microm is determined using FEA The results are plotted in Figure 3-9 and

compared to a structure of constant stiffness ky=035 N m-1 predicted by Eq (15) The

24

non-linearity is greatly reduced when compared to the clamp-clamped flexure with Eq

accurately predicting the deflection until 4 microm However the flexure is more

susceptible to off-axis deflection as there is an order of magnitude reduction in the

stiffness ratio as compared to the clamped-clamped flexure

(15)

The folded beam flexure stands out as it is designed to exhibit a large linear

displacement range and to relieve residual stress due to the manufacturing process by

allowing the beams to expand and contract along the axial direction [40] It was first used

in an electrostatic resonator and has subsequently been used for electrostatic actuation in

a magnetic hard drives and micro-positioning [3344] If the truss section is rigid

ItrussgtgtIbeam then the stiffness of the flexure is expressed as [4043]

3

32b

yEt wk

L=

(17)

The stiffness ratio is expressed as [40]

2x

y

k Lk w

=

(18)

Figure 3-6 Clamped-clamped flexure

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

x03 larger gap distance on the tri-plate comb drive

XN noise displacement y0 initial comb finger overlap length yB base displacement YB base displacement amplitude ym absolute proof mass displacement α parameter dependant on the WplateLplate ratio ∆C Change in capacitance from a tri-plate drive in a differential

arrangement ε permittivity of free space κ second moment of inertia ratio (shinthigh) ζ damping ratio microair absolute viscosity of air ω base excitation frequency in rads ωn sensor natural frequency in rads ωv electrostatic actuation voltage frequency in rads L1 Sensor fabricated using SoiMUMPS with a lateral comb drive T1 Sensor fabricated using PolyMUMPS with a transverse comb drive T2 Sensor fabricated using SoiMUMPS with a tri-plate comb drive

x

1 INTRODUCTION

11 RESEARCH PURPOSE

The modern automobile consists of many mechanical systems such as power seats

windshield wipers mirrors trunks and windows which are all susceptible to breakdown

Without any condition monitoring system the breakdown is usually catastrophic and

requires an expensive part replacement Real-time condition monitoring allows for early

detection of faults which could require a simple solution such as the application of a

lubricant to fix This prolongs the useful life of the component and prevents sudden and

unexpected failure Real-time condition monitoring can be accomplished by examining the

vibration signature of a mechanical system For example an automobile power window

consists of a DC motor and its associated bearings and couplings a gear reduction system

consisting of worm and spur gears and kinematic links Faults resulting in excessive

vibrations may be caused by coupling misalignment bearing failure or gear train failure

Coupling misalignments occur at the connection between the drive shaft and the driven

shaft and are typically due to imperfect manufacturing Bearing failure is caused by lack

of lubrication or moisture contamination causing rusting while gear train failure is caused

by misaligned broken cracked or chipped gear teeth [1] Each fault occurs at a

characteristic frequency and so the state of the mechanical system can be determined by

monitoring the amplitudes of the relevant frequencies Vibrations due to coupling

misalignments occur at harmonics of the shaft rotational speed Gear vibrations occur at

the gear turn speed or at sidebands of the gear mesh frequency [1] Ball bearing vibrations

1

2

may be caused by outer bearing race defects fOB inner bearing race defects fIB ball

defects fB and train defects fT which all occur at specific frequencies [12]

cos( )12

dBOB r

p

bNf fd

β = minus

cos( )12

dBIB r

p

bNf fd

β = +

2cos( )1

2p r d

Bd p

d f bfb d

β = minus

cos( )12

drT

p

bffd

β = minus

(1)

where NB is the number of balls fr is the rotation frequency bd is the ball diameter β is the

contact angle between the ball and races and dp is the ball pitch diameter By monitoring

the real time vibration signature of the mechanical system anomalies can be quickly

identified and fixed

12 MICROELECTROMECHANICAL SENSORS

Microelectromechanical systems (MEMS) inertial sensors provide a small footprint

with sensitivities that are either comparable or exceed any macro sensor along with the

capability of mass production and low unit cost These sensors utilize compliant micro-

flexures attached to a proof mass that displaces in response to an environmental

acceleration Many transduction mechanisms have been developed that convert the

displacement into a measurable electric signal and include thermal piezoresistive

piezoelectric optical and capacitive methods Most MEMS sensors are silicon based and

are fabricated using surface or bulk micromachining Surface micromachining creates free

standing movable structures on top of a substrate using a combination of sacrificial layers

3

and structural layers which are commonly polysilicon [3] In bulk micromachining the

mechanical structures are defined using a removal process where bulk material typically

silicon is etched away MEMS sensors have been used for vibration and shock monitoring

on industrial systems and robotics guidance and navigation in global positioning systems

(GPS) seismometry in earthquake prediction image stabilization in digital cameras and

automobile safety and stability [4]

13 AUTOMOTIVE SENSORS

Sensors cover every major aspect of a modern automobile power-train sensors

monitor fuel combustion and emissions chassis sensors monitor road traction and tire

condition and body sensors facilitate air-bag deployment and vehicle proximity for radar

guided cruise control [5] Pressure sensors which typically consist of a piezoresistive

strain sensing element attached to a silicon diaphragm that deflects when exposed to an

applied pressure are one of the first micro-machined sensors used in an automobile

Implemented as manifold absolute pressure (MAP) sensors they allow precise control of

the air fuel ratio which allows the catalytic converter to efficiently reduce tailpipe

emissions [6] Variable reluctance sensors based on electromagnetics are used for

automobile traction control and produce a voltage output that is dependant on the

magnetic flux variations between a rotating component and the sensor bias magnet [57]

MEMS based linear accelerometers are utilized for airbag deployment upon impact and

provide a lower cost and smaller package over the older ball-in-tube accelerometers [6]

4

14 OPERATING ENVIRONMENT

Through proper packaging MEMS sensors can withstand the harsh automotive

environment where they face cyclic temperatures (-40degC to 150degC) mechanical shock or

vibration (10g-500g) and exposure to various fluids (oil brake fluid or engine coolant)

and airborne particles [8] Temperature changes induce thermal stresses causing

undesirable bending of the silicon microstructures that changes their dynamic response

Mechanical shock which could occur during a high impact collision or vibration from

uneven road conditions cause the delicate microstructures to fracture at points of stress

concentration These effects could be reduced by efficient design Use of a folded beam

micro-flexure that allows expansion and contraction or use of fillets at corners which

reduce the chance of fracture Bonding pads wire bonding and onboard circuit

components can be protected from air borne particles which could cause shorting using

parylene or gels Alternatively the entire package can be sealed hermetically which also

protects the MEMS device [8] In some cases such as fluid pressure sensors intimate

contact between the MEMS device and potentially corrosive fluid is necessary A recent

approach used a nanometer thin coating of silicon carbide (SiC) over a foundry fabricated

MEMS sensor using low temperature low pressure chemical vapor deposition (LPCVD)

[9] The SiC known for its chemical inertness did not compromise the mechanical

characteristics of the original design Therefore MEMS provides a strong platform for an

automotive sensor

5

15 OBJECTIVES AND CONTRIBUTION

The objective of this work is the development of a low frequency MEMS vibration

sensor The sensor must be low cost and mass producible which can be achieved using

standard foundry fabrication processes such as the Polysilicon Multi-User MEMS

Processes (PolyMUMPS) or the Silicon-on-Insulator Multi-User MEMS Processes

(SoiMUMPS) In addition the number of additional processing steps which are required

if integrated CMOS circuitry is used or if vacuum packaging is required must be

minimized This can be achieved using an off-chip circuit made with widely available

discrete components and sensor operation in air

In addition this work presents a MEMS sensor that is fabricated using standard

foundry processes such as PolyMUMPS and SoiMUMPS compared to many MEMS

sensors in literature that are fabricated using customized processes

16 RESEARCH OUTLINE

This work demonstrates three iterations of a low frequency capacitive vibration

sensor fabricated using the standard foundry processes Each sensor consists of a

suspended proof mass that displaces in response to an external vibration The proof mass

is attached to a folded beam micro-flexure compliant in one direction and its lateral

movement is sensed using a capacitive comb drive The first sensor T1 is fabricated in

PolyMUMPS and uses a transverse capacitive comb drive The second L1 fabricated in

SoiMUMPS uses a fully linear lateral capacitive drive and achieves an increase in

sensitivity from T1 Finally the third sensor T2 is fabricated in SoiMUMPS and uses a

differential tri-plate capacitive drive This sensor takes advantage of the increase in

6

sensitivity gained from a transverse drive but rejects off-axis displacement errors and

common mode using a differential tri-plate arrangement In addition an off-chip

capacitance to voltage readout circuit is fabricated and tested on sensor L1

Chapter 2 begins with an overview on the various methods of fault detection

followed by a review of MEMS vibration sensors each with a different transduction

mechanism Chapter 3 outlines the sensor structure focusing on the mechanical design

and sensing principal Chapter 4 focuses on the readout circuitry layout and simulation

results Chapter 5 outlines the various source of noise including mechanical noise arising

from the sensor and electric noise from the readout circuit Chapter 6 overviews the

foundry fabrication processes used including PolyMUMPs and SoiMUMPS The

experimental results of two sensors T1 and L1 which were fabricated and tested within

the thesis time frame are presented in Chapter 7 Chapter 8 concludes the thesis with a

brief summary and outline for future work

2 LITERATURE REVIEW

21 EXISTING METHODS FOR FAULT DETECTION

As this thesis develops a sensor for fault detection using vibration monitoring this

section overviews three existing methods for fault detection in mechanical systems First

motor current monitoring is discussed as it has the ability to measure significant bearing

faults in motor systems Next temperature sensors specifically thin film thermocouples

(TFTC) and MEMS temperature sensors are discussed These sensors are mounted in the

chamber of a lubricating fluid or in close proximity to a moving component and monitor

temperature fluctuations caused by increased wear Finally acoustic emission (AE)

sensors are discussed which have the ability to detect ultrasonic emissions generated by

plastic deformation in the crystalline lattice of a mechanical component

211 MOTOR CURRENT MONITORING

Monitoring of stator current is used as a method of determining bearing related faults

that produce radial motion between rotor and stator This motion varies the air gap flux

density producing stator currents at predictable frequencies fbng that represent the

bearing faults [10]

bng e vf f m f= plusmn sdot (2)

where fe is the electrical supply frequency m=123hellip and fv is one of the characteristics

frequencies listed in the equations of (1)The bearing damage must be significant to

produce noticeable current fluctuations and it may be difficult to distinguish from current

fluctuations due to voltage imbalances machine saturation or statorarmature faults [2]

7

8

212 TEMPERATURE SENSORS

Temperature sensors are largely used in high temperature high load and high speed

environments typically turbine engines and industrial tools Thermocouple sensors

consist of two dissimilar metals joined together at two junctions and when exposed to a

temperature difference generate a thermoelectric voltage dependant only on the material

properties and junction temperature difference [11] In turbine engines these sensors can

be placed on bearing casings or in the lubricant to monitor temperature fluctuations due

to increased friction The thermocouples are fabricated using bulk materials or micro-

machined thin films (TFTC) with the latter providing a faster response small package

and substantially lower production costs [12] The TFTC can be embedded for in situ

monitoring of processes in hostile environments One study successfully embedded

TFTC into electroplated nickel and attained device characteristics on par with surface

mounted devices [13] More recently MEMS temperature sensors have been developed

that consist of micro-machined semiconductor material that undergoes structural

deformation with temperature changes A common implementation is a bent beam

structure whose temperature induced deflection δ is expressed as [14]

( ) 3

2 2

sin12 cos sinz

EA T LEI AEL

α βδ 2β β

∆=

+

(3)

Temperature sensors are well suited for localized fault monitoring and detection but can

only characterize a mechanical system consisting of multiple components through the use

of multiple sensors

9

213 ACOUSTIC EMISSION SENSORS

Acoustic emission (AE) sensors have been used to characterize wear in machine

tools and monitor bearing and gear problems in centrifugal pumps [1516] First

developed as a Non-Destructive Testing (NDT) technique to detect cracks in civil

structures these sensors detect acoustic emissions generated by the release of vibration

waves in a crystalline lattice due to plastic deformation or crack growth [16]

Measurements are made using piezoelectric transducers with high natural frequencies

100 kHz to 1 MHz to capture the ultrasonic AE emissions An AE sensor is useful as it

has the ability to detect subsurface cracks in gear teeth or bearings before appearing on

the surface causing further damage [17] More recently MEMS acoustic sensors have

been developed and one design includes multiple transducers on a single substrate

which each detect acoustic emission energy at different frequencies [18] This helps

distinguish spurious acoustic emissions arising from impact and friction from those

arising from plastic deformation When compared to typical piezoelectric sensors the

MEMS devices have lower sensitivities and fail to detect some acoustic emissions

[1819] In addition the acoustic emission signal suffers severe attenuation as is crosses

various interfaces such as a gearbox or bearing casings In one experiment consisting of

a pinion gear and an associated bearing a 44dB attenuation was seen between an AE

sensor placed directly on the pinion to one placed on the bearing casing and in some

cases intermediate loss of the signal was observed [20]

10

22 MEMS VIBRATION SENSORS

This thesis focuses on the development of a MEMS vibration sensor for fault

detection as the vibration signature of a mechanical system has the potential to give an

overview of the entire system Many types of MEMS vibration sensors exist with

different operating principals based on resonance the piezoelectric effect and

displacement variation

221 RESONANT SENSORS

Resonant sensors are based on the resonant frequency shift of a fixed-fixed micro-

beam that is excited into resonance using electrostatic thermal or piezoelectric methods

Under an applied axial load there is a shift in the resonant frequency which is expressed

as [21]

2 2

2

1 12 1

nn n

EI Fl2

fl A Eα γ

π ρ= +

I

(4)

where l is the beam length E is Youngrsquos Modulus I is the moment of inertia ρ is the

density A is the cross section area F is the applied axial force αn and γn are mode

dependant coefficients and are 4730 and 0295 for mode 1 in the case discussed The

axial force F is generated by a proof mass attached to one beam end which moves in

response to an external acceleration The performance of resonant sensors are dependant

on the quality factor Q which is the ratio of total energy stored in the system to the

energy lost per cycle and determines the sharpness and amplitude of the resonance peak

A high Q resonator develops a peak that is easily distinguishable in a phase locked loop

control circuit and allows for improved performance and resolution [3] A resonator Q

value is lowered by energy losses to the surrounding fluid attached supports and

11

material Losses to the fluid surrounding the MEMS structure is the dominant loss

mechanism for MEMS sensors operated in air and can be greatly reduced by vacuum

packaging which can be accomplished using glass-silicon anodic bonding [22] In

addition to increasing fabrication cost the long term stability of the vacuum degrades

with ultimate device failure stemming from leakage through micro cracks and defects

and out-gassing [22] Support losses arise from the restoring forces they generate and

can be reduce by balanced designs such as a double ended tuning fork (DETF) or

operation in higher modes [3] In addition the sensitivity of a resonant sensor can be

increased by the addition of micro-levers to increase the axial force In one such design a

two stage micro-lever increased the sensitivity by an order of magnitude over a more

conventional single lever arm [2324]

222 PIEZOELECTRIC SENSORS

Piezoelectric materials produce a charge in response to an applied force This is an

intrinsic effect in materials such as Zinc Oxide (ZnO) whereas materials such as Barium

Titanate or Lead Zirconate Titanate (PZT) need to be poled by placing the material in a

strong electric field at an elevated temperature [3] The piezoelectric effect is expressed

mathematical by [25]

i ij j ik kD d T Eε= + (5)

where D represents the electrical displacement (Cm2) d is the charge coefficient (CN)

T is the applied stress matrix (Nm2) ε is the permittivity matrix (Fm) and E is the

electric field matrix (Vm) The subscripts i=123 j=123456 and k=123 represent

the direction to which the physical property is related For micro-sensors poled in the out

of plane direction only the D3 term is relevant and if fabricated using ZnO or PZT only

12

the charge coefficients d31 d32 and d33 are relevant [25] The piezoelectric material is

commonly deposited as a thin film on the surface of compliant structures in bulk micro-

machined sensors or incorporated as a layer in surface micro-machining on a similar

structure [26] However the inclusion of multiple materials introduces large temperature

variations causing out of plane bending especially in surface micro-machined devices

[2627] Poled piezoelectric materials could be depolarized which serves to reduce the

piezoelectric affect in the material This can occur if the sensor is exposed to a strong

electric field of the opposite polarity high temperatures in excess of the Curie point or

high mechanical stress Also the piezoelectric effect reduces over time an effect that can

be reduced with the addition of composite elements [28] Foundry processes such as

PolyMUMPS and SoiMUMPS do not incorporate piezoelectric layers thus sensors using

this method of transduction must use a customized process The piezoelectric sensor

outlined in [26] uses polysilicon surface micro-machining similar to PolyMUMPS

however it incorporates a ZnO layer that is not available in the foundry process

223 DISPLACEMENT VARIATION SENSORS

Displacement variation sensors consist of a proof mass connected to a compliant

micro-flexure The movement of the proof mass is sensed using optical electron

tunneling or more commonly capacitive methods A recent optical accelerometer

utilized a novel multilayer nano-grating for in-plane displacement sensing of the proof

mass [29] For small space variations between the nano-gratings a large change in the

optical reflection amplitude of the grating was observed [30] Even though optical

sensors offer resolutions approaching the Brownian noise limit they are not available in

small packages that are easily placed in space constricted areas for a low cost [29]

13

The electron tunneling effect is observed when the proximity between two electrodes

one flexible and the other fixed is on the order of 10Ǻ [31] If the two electrodes are

biased with voltage VB a tunneling current It flows between the gap [31]

exp( )t B I tgI V xαinfin sdot minus Φ (6)

where α1 is a constant and equal to 1025 Ǻ-1 eV-frac12 Ф is the effective height of the

tunneling barrier and xtg is the tunneling gap When the flexible electrode moves in

response to acceleration a feedback controller maintains the original distance while

determining the applied acceleration Electron tunneling sensors are able to sense

accelerations in the microg-ng range due to the exponential relation between tunneling current

and gap However these highly sensitive devices are susceptible to thermo-mechanical

noise which could be reduced by operation at low pressure [32] It is also important for

the two electrodes to be coated with metal which is difficult to achieve with foundry

fabrication

Capacitive sensors are based on the parallel plate capacitor and are implemented

using inter-digitated comb fingers in an in-plane lateral or transverse drive or in an out of

plane parallel plate drive [33] These sensors have the advantage of high sensitivity low

power consumption small package low temperature dependence and are easily

integrated with existing foundry fabrication processes [34] The lateral drive offers a fully

linear response but low sensitivity whereas the transverse and parallel plate drives

achieve high sensitivity if gap distances are of a few microns and large capacitance areas

are utilized However both drives have a non-linear response and are prone to pull-in

Also if these sensors are fabricated using a surface micromachining process such as

PolyMUMPS where the ground plane is a few microns above the structural layer they

14

are prone to electrical failure due to shorting This was observed in an experiment where

similar sensors were placed in a vibrating environment with a peak of 120g and shorting

between the structural layer and ground plane was found to be the predominate mode of

failure [35] However with certain fabrication processes such as SoiMUMPS this issue

can be neglected as the substrate material below any moving component is removed

Most capacitance sensors are fabricated using an integrated CMOS (Complementary

Metal Oxide Semiconductor)-MEMS technique where the MEMS sensor is fabricated

before during or after the CMOS circuit fabrication [36] Although these sensors offer

high-sensitivity and small parasitic capacitance they suffer from in-plane and out of

plane curling of the beam which reduces the capacitance between adjacent comb fingers

[37] In addition multiple processing steps are required for device fabrication

More than the other transduction methods MEMS capacitive displacement sensors are

attractive for practical implementation in an automobile and therefore is the selected

method for the vibration sensor presented in this work Their low temperature dependence

is an ideal characteristic for an automobile environment that faces varying temperature If

implemented with off-chip readout circuitry they require a standard foundry fabrication

process which provides a strong platform for mass production and low unit cost

3 SENSOR STRUCTURE

This chapter presents the physical structure of the three MEMS vibration sensors

which includes the displacement sensing capacitive comb drive and the micro-

flexureproof mass system First the capacitive comb drive and the three variations

implemented the transverse drive lateral drive and differential tri-plate drive are

discussed Next the mechanical structure is presented and includes the micro-flexure

selection quality factor and equations of motion Finally instability that places a limit

on the operating range of the sensor as a result of a gap closing capacitive comb drive is

discussed

31 SENSING

Capacitive sensors are based on the parallel plate capacitor and are commonly

implemented using inter-digitated comb fingers A capacitance can be realized if a voltage

is placed between two closely spaced plates If the fringing electric field is neglected the

capacitance of the parallel plate capacitor Cpp (F) is expressed as [28]

air pppp

k AC

dε

= (7)

where kair=1 is the dielectric constant of air at atmospheric pressure ε=88542times10-12 F m-1

is the permittivity of free space App is the overlapping area (m2) and d is the distance

between the two plates (m)

15

16

Figure 3-1 An inter-digitated comb drive where changes in capacitance are generated by either changes in gap distance x0 or in the overlap area y0timest

Figure 3-1

For a set of inter-digitated comb fingers or comb drive shown in the total

capacitance Ctotal is expressed as [28]

0 0

0 0

( ) ( )( ) (Total cy y y yC N t

)x x x xε

+ += + + minus

(8)

where N is the number of comb fingers on the movable comb y0 is the initial comb finger

overlap length (m) tc is the out of plane comb finger thickness (m) and x0 is the inital

finger gap (m) In this work the comb drive is implemented in three arrangements

transverse lateral and differential tri-plate In a transverse comb drive changes in

capacitance are generated by transverse movements along the x-axis resulting in changes

to the gap x0 between the movable and fixed comb fingers The nonlinear relation

between total capacitance CTotalT and displacement is expressed as

17

00 0

1 1( ) (Total TC N ty

)x x x xε

= + minus +

(9)

Although the drive sensitivity part is not constant an average sensitivity can be

assumed for small displacements [38]

Total TC xpart

In a lateral comb drive changes in capacitance are generated by lateral movements

along the y-axis resulting in changes to the overlapping area y0timestc The total capacitance

CTotalL between all comb fingers is expressed as

00

2 ( )Total LN tC yx

yε= +

(10)

The sensitivity of the lateral comb drive SL is expressed as

0

2Total LL

C N tSy x

εpart= =

part

(11)

The lateral comb drive was first introduced by W C Tang T H Nguyen and R T

Howe [33] in an electrostatic resonator and has since been used numerous times in both

drive and sense configurations It has the advantage of a fully linear response and

negligible instability but offers low sensitivity The sensitivity for the lateral drive L1

implemented in this work is 0038 pF microm-1

A differential tri-plate drive shown in Figure 3-2 which is a variation of the

transverse drive is implemented to improve sensitivity and reduce noise [39]

18

Figure 3-2 An inter-digitated comb drive in a differential tri-plate arrangement where x01 x02 and x03 are the gap distances and x01= x02ltlt x03 to improve sensitivity

In the tri-plate drive the at rest gap distances between the two capacitors formed by any

movable finger is not equivalent For a differential arrangement insuring that x01=

x02ltltx03 the linearity for small displacements (less than 1 microm) is enhanced over the case

of a single ended arrangement where x01=x02=x03=x0 which is equivalent to the

transverse drive This is shown graphically in Figure 3-3 where x01= x02=x03=2 microm in

the lower curve and x01= x02= 2 microm and x03= 15 microm in the upper curve A substantial

improvement in linearity is attained with the tri-plate drive

19

0

005

01

015

02

025

03

035

0 02 04 06 08 1

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-3 Comparison of the effect of gap distances on linearity for two case lsquorsquo represents case 1 where the fingers in are implemented in a differential tri-plate arrangement and x01= x02= 2 microm and x03= 15 microm lsquorsquo represents the case 2 where the fingers are implemented in a single ended arrangement and x01= x02=x03=2 microm

The tri-plate drive is also implemented in a differential arrangement where the output

is the change in capacitance rather than the total capacitance resulting from an external

excitation This reduces errors cause by off-axis excitation and common mode noise The

output of the differential tri-plate drive is expressed as [39]

Total Top Total BottomC C C∆ = minus where

003 01

1 1( ) (Total TopC N ty

)x x x xε

= +

minus +

(12)

20

002 03

1 1( ) (Total BottomC N ty

)x x x xε

= +

minus +

A graphic comparison of the capacitancedisplacement relation for sensors T1 and T2 is

shown in Figure 3-4 For these gap closing drives it is assumed that for displacements

less that 1 microm the relation is linear indicated by the dashed line above each curve The

sensitivity within this region is approximately 0046 pF microm-1 and 0253 pF microm-1 for T1

and T2 respectively The order of magnitude improvement is attributed to the differential

tri-plate arrangement and thicker comb finger

0

02

04

06

08

1

12

14

16

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-4 lsquorsquo represents the output of the differential tri-plate drive while lsquorsquo represents the output of the transverse drive The dashed line represents the linear approximation for displacements less that 1 microm For the purpose of comparison the output of the transverse drive was equalized to zero at zero displacement

21

In this work the transverse drive is first implemented in sensor T1 using the

PolyMUMPS fabrication process Then in an effort to linearize the sensor response the

lateral drive is used in L1 using SoiMUMPS The reduction in drive sensitivity due to the

lateral drive is further reduced by increasing the gap distance to 3 microm which is

recommended by the foundry to ensure a successful device These reductions are

balanced by increasing the number of comb fingers N and selecting a substrate thickness

that is 125 times large than T1 The third device T2 is fabricated using SoiMUMPS and

uses the tri-plate differential drive This design capitalizes on the increase in sensitivity of

a gap closing drive and larger substrate thickness attaining an order of magnitude

increase in sensitivity over T1 and L1 The specifications for each drive along with the

successive improvement in drive sensitivity are outlined in Table 3-1

Table 3-1 Specifications for each capacitive comb drive implemented in this work T1 L1 and T2

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Comb Finger Thickness tc (microm) 2 25 25 Gap Distance x0 (microm) 2 3 2 Number of Fingers N 200 256 30 Overlap length y0 (microm) 45 10 70 Drive Sensitivity (pF microm-1) 0046 0038 0273

32 MECHANICAL ANALYSIS

321 MICRO-FLEXURE SELECTION

The sensor structure consists of a proof mass connected to a compliant micro-flexure

that itself is anchored to the sensor substrate The proof mass displaces in response to an

22

applied acceleration and this displacement is sensed using the capacitive comb drive

attached to the proof mass

Proof mass

Folded beam micro-flexure

Anchor

Figure 3-5 Sensor Structure

This section focuses on micro-flexure selection as it should exhibit a large linear

displacement range and have a high lateral stiffness ratio to reduce cross-axis sensitivity

A number of micro-flexures have been used in MEMS research these include the

clamped-clamped crab-leg and folded beam shown in Figure 3-6-through 3-7

The clamped-clamped flexure exhibits a high stiffness ratio however for large lateral

displacements extensional axial forces buildup within the beams resulting in a non-linear

force-displacement relation [40] Within the linear regime the stiffness of the clamped-

clamped flexure is expressed as [40]

3

3b

yEt wk

L=

(13)

23

where E is the Youngrsquos Modulus of the beam material (N m-2) tb is the beam out of

plane thickness (m) w is the beam width (m) and L is the beam length (m) The stiffness

ratio is expressed as [40]

2x

y

k Lk w

=

(14)

If L is two orders of magnitude greater than w this results in a ratio of approximately

10000 The linear displacement range is determined using Finite Element Analysis (FEA)

of a clamped-clamped flexure with L=387 microm w=4 microm and tb=2 microm The result of this

simulation is plotted in and compared to a structure with constant stiffness ky

= 035 N m-1 predicted by Eq (13) The linear relation is valid for small deflections that

are less than 2 microm In addition this micro-flexure is susceptible to buckling as a result of

a compressive residual stress that may arise due to the fabrication process [41]

Figure 3-9

In the crab leg design the thigh section is meant to relieve the extensional axial

stress postponing the occurrence of non-linearity until greater lateral displacement [42]

The stiffness of the crab leg flexure is expressed as [43]

3

44

b s s ty

s s t

Et w L LkL L L

κκ

+= +

(15)

where κ=IsIt=(wswt)3 is a ratio of the second moment of inertia of the shin (s) and thigh

(t) ws and wt represent the width and Ls and Lt represent the length of the shin and thigh

regions respectively The stiffness ratio is expressed as [43]

34

4t s s tx

y s t s

w L L Lkk w L L L

αα t

+= +

(16)

The linear displacement range of a crab leg flexure with Ls=285 microm Lt=30 microm ws=wt=3

microm and tb=2 microm is determined using FEA The results are plotted in Figure 3-9 and

compared to a structure of constant stiffness ky=035 N m-1 predicted by Eq (15) The

24

non-linearity is greatly reduced when compared to the clamp-clamped flexure with Eq

accurately predicting the deflection until 4 microm However the flexure is more

susceptible to off-axis deflection as there is an order of magnitude reduction in the

stiffness ratio as compared to the clamped-clamped flexure

(15)

The folded beam flexure stands out as it is designed to exhibit a large linear

displacement range and to relieve residual stress due to the manufacturing process by

allowing the beams to expand and contract along the axial direction [40] It was first used

in an electrostatic resonator and has subsequently been used for electrostatic actuation in

a magnetic hard drives and micro-positioning [3344] If the truss section is rigid

ItrussgtgtIbeam then the stiffness of the flexure is expressed as [4043]

3

32b

yEt wk

L=

(17)

The stiffness ratio is expressed as [40]

2x

y

k Lk w

=

(18)

Figure 3-6 Clamped-clamped flexure

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

1 INTRODUCTION

11 RESEARCH PURPOSE

The modern automobile consists of many mechanical systems such as power seats

windshield wipers mirrors trunks and windows which are all susceptible to breakdown

Without any condition monitoring system the breakdown is usually catastrophic and

requires an expensive part replacement Real-time condition monitoring allows for early

detection of faults which could require a simple solution such as the application of a

lubricant to fix This prolongs the useful life of the component and prevents sudden and

unexpected failure Real-time condition monitoring can be accomplished by examining the

vibration signature of a mechanical system For example an automobile power window

consists of a DC motor and its associated bearings and couplings a gear reduction system

consisting of worm and spur gears and kinematic links Faults resulting in excessive

vibrations may be caused by coupling misalignment bearing failure or gear train failure

Coupling misalignments occur at the connection between the drive shaft and the driven

shaft and are typically due to imperfect manufacturing Bearing failure is caused by lack

of lubrication or moisture contamination causing rusting while gear train failure is caused

by misaligned broken cracked or chipped gear teeth [1] Each fault occurs at a

characteristic frequency and so the state of the mechanical system can be determined by

monitoring the amplitudes of the relevant frequencies Vibrations due to coupling

misalignments occur at harmonics of the shaft rotational speed Gear vibrations occur at

the gear turn speed or at sidebands of the gear mesh frequency [1] Ball bearing vibrations

1

2

may be caused by outer bearing race defects fOB inner bearing race defects fIB ball

defects fB and train defects fT which all occur at specific frequencies [12]

cos( )12

dBOB r

p

bNf fd

β = minus

cos( )12

dBIB r

p

bNf fd

β = +

2cos( )1

2p r d

Bd p

d f bfb d

β = minus

cos( )12

drT

p

bffd

β = minus

(1)

where NB is the number of balls fr is the rotation frequency bd is the ball diameter β is the

contact angle between the ball and races and dp is the ball pitch diameter By monitoring

the real time vibration signature of the mechanical system anomalies can be quickly

identified and fixed

12 MICROELECTROMECHANICAL SENSORS

Microelectromechanical systems (MEMS) inertial sensors provide a small footprint

with sensitivities that are either comparable or exceed any macro sensor along with the

capability of mass production and low unit cost These sensors utilize compliant micro-

flexures attached to a proof mass that displaces in response to an environmental

acceleration Many transduction mechanisms have been developed that convert the

displacement into a measurable electric signal and include thermal piezoresistive

piezoelectric optical and capacitive methods Most MEMS sensors are silicon based and

are fabricated using surface or bulk micromachining Surface micromachining creates free

standing movable structures on top of a substrate using a combination of sacrificial layers

3

and structural layers which are commonly polysilicon [3] In bulk micromachining the

mechanical structures are defined using a removal process where bulk material typically

silicon is etched away MEMS sensors have been used for vibration and shock monitoring

on industrial systems and robotics guidance and navigation in global positioning systems

(GPS) seismometry in earthquake prediction image stabilization in digital cameras and

automobile safety and stability [4]

13 AUTOMOTIVE SENSORS

Sensors cover every major aspect of a modern automobile power-train sensors

monitor fuel combustion and emissions chassis sensors monitor road traction and tire

condition and body sensors facilitate air-bag deployment and vehicle proximity for radar

guided cruise control [5] Pressure sensors which typically consist of a piezoresistive

strain sensing element attached to a silicon diaphragm that deflects when exposed to an

applied pressure are one of the first micro-machined sensors used in an automobile

Implemented as manifold absolute pressure (MAP) sensors they allow precise control of

the air fuel ratio which allows the catalytic converter to efficiently reduce tailpipe

emissions [6] Variable reluctance sensors based on electromagnetics are used for

automobile traction control and produce a voltage output that is dependant on the

magnetic flux variations between a rotating component and the sensor bias magnet [57]

MEMS based linear accelerometers are utilized for airbag deployment upon impact and

provide a lower cost and smaller package over the older ball-in-tube accelerometers [6]

4

14 OPERATING ENVIRONMENT

Through proper packaging MEMS sensors can withstand the harsh automotive

environment where they face cyclic temperatures (-40degC to 150degC) mechanical shock or

vibration (10g-500g) and exposure to various fluids (oil brake fluid or engine coolant)

and airborne particles [8] Temperature changes induce thermal stresses causing

undesirable bending of the silicon microstructures that changes their dynamic response

Mechanical shock which could occur during a high impact collision or vibration from

uneven road conditions cause the delicate microstructures to fracture at points of stress

concentration These effects could be reduced by efficient design Use of a folded beam

micro-flexure that allows expansion and contraction or use of fillets at corners which

reduce the chance of fracture Bonding pads wire bonding and onboard circuit

components can be protected from air borne particles which could cause shorting using

parylene or gels Alternatively the entire package can be sealed hermetically which also

protects the MEMS device [8] In some cases such as fluid pressure sensors intimate

contact between the MEMS device and potentially corrosive fluid is necessary A recent

approach used a nanometer thin coating of silicon carbide (SiC) over a foundry fabricated

MEMS sensor using low temperature low pressure chemical vapor deposition (LPCVD)

[9] The SiC known for its chemical inertness did not compromise the mechanical

characteristics of the original design Therefore MEMS provides a strong platform for an

automotive sensor

5

15 OBJECTIVES AND CONTRIBUTION

The objective of this work is the development of a low frequency MEMS vibration

sensor The sensor must be low cost and mass producible which can be achieved using

standard foundry fabrication processes such as the Polysilicon Multi-User MEMS

Processes (PolyMUMPS) or the Silicon-on-Insulator Multi-User MEMS Processes

(SoiMUMPS) In addition the number of additional processing steps which are required

if integrated CMOS circuitry is used or if vacuum packaging is required must be

minimized This can be achieved using an off-chip circuit made with widely available

discrete components and sensor operation in air

In addition this work presents a MEMS sensor that is fabricated using standard

foundry processes such as PolyMUMPS and SoiMUMPS compared to many MEMS

sensors in literature that are fabricated using customized processes

16 RESEARCH OUTLINE

This work demonstrates three iterations of a low frequency capacitive vibration

sensor fabricated using the standard foundry processes Each sensor consists of a

suspended proof mass that displaces in response to an external vibration The proof mass

is attached to a folded beam micro-flexure compliant in one direction and its lateral

movement is sensed using a capacitive comb drive The first sensor T1 is fabricated in

PolyMUMPS and uses a transverse capacitive comb drive The second L1 fabricated in

SoiMUMPS uses a fully linear lateral capacitive drive and achieves an increase in

sensitivity from T1 Finally the third sensor T2 is fabricated in SoiMUMPS and uses a

differential tri-plate capacitive drive This sensor takes advantage of the increase in

6

sensitivity gained from a transverse drive but rejects off-axis displacement errors and

common mode using a differential tri-plate arrangement In addition an off-chip

capacitance to voltage readout circuit is fabricated and tested on sensor L1

Chapter 2 begins with an overview on the various methods of fault detection

followed by a review of MEMS vibration sensors each with a different transduction

mechanism Chapter 3 outlines the sensor structure focusing on the mechanical design

and sensing principal Chapter 4 focuses on the readout circuitry layout and simulation

results Chapter 5 outlines the various source of noise including mechanical noise arising

from the sensor and electric noise from the readout circuit Chapter 6 overviews the

foundry fabrication processes used including PolyMUMPs and SoiMUMPS The

experimental results of two sensors T1 and L1 which were fabricated and tested within

the thesis time frame are presented in Chapter 7 Chapter 8 concludes the thesis with a

brief summary and outline for future work

2 LITERATURE REVIEW

21 EXISTING METHODS FOR FAULT DETECTION

As this thesis develops a sensor for fault detection using vibration monitoring this

section overviews three existing methods for fault detection in mechanical systems First

motor current monitoring is discussed as it has the ability to measure significant bearing

faults in motor systems Next temperature sensors specifically thin film thermocouples

(TFTC) and MEMS temperature sensors are discussed These sensors are mounted in the

chamber of a lubricating fluid or in close proximity to a moving component and monitor

temperature fluctuations caused by increased wear Finally acoustic emission (AE)

sensors are discussed which have the ability to detect ultrasonic emissions generated by

plastic deformation in the crystalline lattice of a mechanical component

211 MOTOR CURRENT MONITORING

Monitoring of stator current is used as a method of determining bearing related faults

that produce radial motion between rotor and stator This motion varies the air gap flux

density producing stator currents at predictable frequencies fbng that represent the

bearing faults [10]

bng e vf f m f= plusmn sdot (2)

where fe is the electrical supply frequency m=123hellip and fv is one of the characteristics

frequencies listed in the equations of (1)The bearing damage must be significant to

produce noticeable current fluctuations and it may be difficult to distinguish from current

fluctuations due to voltage imbalances machine saturation or statorarmature faults [2]

7

8

212 TEMPERATURE SENSORS

Temperature sensors are largely used in high temperature high load and high speed

environments typically turbine engines and industrial tools Thermocouple sensors

consist of two dissimilar metals joined together at two junctions and when exposed to a

temperature difference generate a thermoelectric voltage dependant only on the material

properties and junction temperature difference [11] In turbine engines these sensors can

be placed on bearing casings or in the lubricant to monitor temperature fluctuations due

to increased friction The thermocouples are fabricated using bulk materials or micro-

machined thin films (TFTC) with the latter providing a faster response small package

and substantially lower production costs [12] The TFTC can be embedded for in situ

monitoring of processes in hostile environments One study successfully embedded

TFTC into electroplated nickel and attained device characteristics on par with surface

mounted devices [13] More recently MEMS temperature sensors have been developed

that consist of micro-machined semiconductor material that undergoes structural

deformation with temperature changes A common implementation is a bent beam

structure whose temperature induced deflection δ is expressed as [14]

( ) 3

2 2

sin12 cos sinz

EA T LEI AEL

α βδ 2β β

∆=

+

(3)

Temperature sensors are well suited for localized fault monitoring and detection but can

only characterize a mechanical system consisting of multiple components through the use

of multiple sensors

9

213 ACOUSTIC EMISSION SENSORS

Acoustic emission (AE) sensors have been used to characterize wear in machine

tools and monitor bearing and gear problems in centrifugal pumps [1516] First

developed as a Non-Destructive Testing (NDT) technique to detect cracks in civil

structures these sensors detect acoustic emissions generated by the release of vibration

waves in a crystalline lattice due to plastic deformation or crack growth [16]

Measurements are made using piezoelectric transducers with high natural frequencies

100 kHz to 1 MHz to capture the ultrasonic AE emissions An AE sensor is useful as it

has the ability to detect subsurface cracks in gear teeth or bearings before appearing on

the surface causing further damage [17] More recently MEMS acoustic sensors have

been developed and one design includes multiple transducers on a single substrate

which each detect acoustic emission energy at different frequencies [18] This helps

distinguish spurious acoustic emissions arising from impact and friction from those

arising from plastic deformation When compared to typical piezoelectric sensors the

MEMS devices have lower sensitivities and fail to detect some acoustic emissions

[1819] In addition the acoustic emission signal suffers severe attenuation as is crosses

various interfaces such as a gearbox or bearing casings In one experiment consisting of

a pinion gear and an associated bearing a 44dB attenuation was seen between an AE

sensor placed directly on the pinion to one placed on the bearing casing and in some

cases intermediate loss of the signal was observed [20]

10

22 MEMS VIBRATION SENSORS

This thesis focuses on the development of a MEMS vibration sensor for fault

detection as the vibration signature of a mechanical system has the potential to give an

overview of the entire system Many types of MEMS vibration sensors exist with

different operating principals based on resonance the piezoelectric effect and

displacement variation

221 RESONANT SENSORS

Resonant sensors are based on the resonant frequency shift of a fixed-fixed micro-

beam that is excited into resonance using electrostatic thermal or piezoelectric methods

Under an applied axial load there is a shift in the resonant frequency which is expressed

as [21]

2 2

2

1 12 1

nn n

EI Fl2

fl A Eα γ

π ρ= +

I

(4)

where l is the beam length E is Youngrsquos Modulus I is the moment of inertia ρ is the

density A is the cross section area F is the applied axial force αn and γn are mode

dependant coefficients and are 4730 and 0295 for mode 1 in the case discussed The

axial force F is generated by a proof mass attached to one beam end which moves in

response to an external acceleration The performance of resonant sensors are dependant

on the quality factor Q which is the ratio of total energy stored in the system to the

energy lost per cycle and determines the sharpness and amplitude of the resonance peak

A high Q resonator develops a peak that is easily distinguishable in a phase locked loop

control circuit and allows for improved performance and resolution [3] A resonator Q

value is lowered by energy losses to the surrounding fluid attached supports and

11

material Losses to the fluid surrounding the MEMS structure is the dominant loss

mechanism for MEMS sensors operated in air and can be greatly reduced by vacuum

packaging which can be accomplished using glass-silicon anodic bonding [22] In

addition to increasing fabrication cost the long term stability of the vacuum degrades

with ultimate device failure stemming from leakage through micro cracks and defects

and out-gassing [22] Support losses arise from the restoring forces they generate and

can be reduce by balanced designs such as a double ended tuning fork (DETF) or

operation in higher modes [3] In addition the sensitivity of a resonant sensor can be

increased by the addition of micro-levers to increase the axial force In one such design a

two stage micro-lever increased the sensitivity by an order of magnitude over a more

conventional single lever arm [2324]

222 PIEZOELECTRIC SENSORS

Piezoelectric materials produce a charge in response to an applied force This is an

intrinsic effect in materials such as Zinc Oxide (ZnO) whereas materials such as Barium

Titanate or Lead Zirconate Titanate (PZT) need to be poled by placing the material in a

strong electric field at an elevated temperature [3] The piezoelectric effect is expressed

mathematical by [25]

i ij j ik kD d T Eε= + (5)

where D represents the electrical displacement (Cm2) d is the charge coefficient (CN)

T is the applied stress matrix (Nm2) ε is the permittivity matrix (Fm) and E is the

electric field matrix (Vm) The subscripts i=123 j=123456 and k=123 represent

the direction to which the physical property is related For micro-sensors poled in the out

of plane direction only the D3 term is relevant and if fabricated using ZnO or PZT only

12

the charge coefficients d31 d32 and d33 are relevant [25] The piezoelectric material is

commonly deposited as a thin film on the surface of compliant structures in bulk micro-

machined sensors or incorporated as a layer in surface micro-machining on a similar

structure [26] However the inclusion of multiple materials introduces large temperature

variations causing out of plane bending especially in surface micro-machined devices

[2627] Poled piezoelectric materials could be depolarized which serves to reduce the

piezoelectric affect in the material This can occur if the sensor is exposed to a strong

electric field of the opposite polarity high temperatures in excess of the Curie point or

high mechanical stress Also the piezoelectric effect reduces over time an effect that can

be reduced with the addition of composite elements [28] Foundry processes such as

PolyMUMPS and SoiMUMPS do not incorporate piezoelectric layers thus sensors using

this method of transduction must use a customized process The piezoelectric sensor

outlined in [26] uses polysilicon surface micro-machining similar to PolyMUMPS

however it incorporates a ZnO layer that is not available in the foundry process

223 DISPLACEMENT VARIATION SENSORS

Displacement variation sensors consist of a proof mass connected to a compliant

micro-flexure The movement of the proof mass is sensed using optical electron

tunneling or more commonly capacitive methods A recent optical accelerometer

utilized a novel multilayer nano-grating for in-plane displacement sensing of the proof

mass [29] For small space variations between the nano-gratings a large change in the

optical reflection amplitude of the grating was observed [30] Even though optical

sensors offer resolutions approaching the Brownian noise limit they are not available in

small packages that are easily placed in space constricted areas for a low cost [29]

13

The electron tunneling effect is observed when the proximity between two electrodes

one flexible and the other fixed is on the order of 10Ǻ [31] If the two electrodes are

biased with voltage VB a tunneling current It flows between the gap [31]

exp( )t B I tgI V xαinfin sdot minus Φ (6)

where α1 is a constant and equal to 1025 Ǻ-1 eV-frac12 Ф is the effective height of the

tunneling barrier and xtg is the tunneling gap When the flexible electrode moves in

response to acceleration a feedback controller maintains the original distance while

determining the applied acceleration Electron tunneling sensors are able to sense

accelerations in the microg-ng range due to the exponential relation between tunneling current

and gap However these highly sensitive devices are susceptible to thermo-mechanical

noise which could be reduced by operation at low pressure [32] It is also important for

the two electrodes to be coated with metal which is difficult to achieve with foundry

fabrication

Capacitive sensors are based on the parallel plate capacitor and are implemented

using inter-digitated comb fingers in an in-plane lateral or transverse drive or in an out of

plane parallel plate drive [33] These sensors have the advantage of high sensitivity low

power consumption small package low temperature dependence and are easily

integrated with existing foundry fabrication processes [34] The lateral drive offers a fully

linear response but low sensitivity whereas the transverse and parallel plate drives

achieve high sensitivity if gap distances are of a few microns and large capacitance areas

are utilized However both drives have a non-linear response and are prone to pull-in

Also if these sensors are fabricated using a surface micromachining process such as

PolyMUMPS where the ground plane is a few microns above the structural layer they

14

are prone to electrical failure due to shorting This was observed in an experiment where

similar sensors were placed in a vibrating environment with a peak of 120g and shorting

between the structural layer and ground plane was found to be the predominate mode of

failure [35] However with certain fabrication processes such as SoiMUMPS this issue

can be neglected as the substrate material below any moving component is removed

Most capacitance sensors are fabricated using an integrated CMOS (Complementary

Metal Oxide Semiconductor)-MEMS technique where the MEMS sensor is fabricated

before during or after the CMOS circuit fabrication [36] Although these sensors offer

high-sensitivity and small parasitic capacitance they suffer from in-plane and out of

plane curling of the beam which reduces the capacitance between adjacent comb fingers

[37] In addition multiple processing steps are required for device fabrication

More than the other transduction methods MEMS capacitive displacement sensors are

attractive for practical implementation in an automobile and therefore is the selected

method for the vibration sensor presented in this work Their low temperature dependence

is an ideal characteristic for an automobile environment that faces varying temperature If

implemented with off-chip readout circuitry they require a standard foundry fabrication

process which provides a strong platform for mass production and low unit cost

3 SENSOR STRUCTURE

This chapter presents the physical structure of the three MEMS vibration sensors

which includes the displacement sensing capacitive comb drive and the micro-

flexureproof mass system First the capacitive comb drive and the three variations

implemented the transverse drive lateral drive and differential tri-plate drive are

discussed Next the mechanical structure is presented and includes the micro-flexure

selection quality factor and equations of motion Finally instability that places a limit

on the operating range of the sensor as a result of a gap closing capacitive comb drive is

discussed

31 SENSING

Capacitive sensors are based on the parallel plate capacitor and are commonly

implemented using inter-digitated comb fingers A capacitance can be realized if a voltage

is placed between two closely spaced plates If the fringing electric field is neglected the

capacitance of the parallel plate capacitor Cpp (F) is expressed as [28]

air pppp

k AC

dε

= (7)

where kair=1 is the dielectric constant of air at atmospheric pressure ε=88542times10-12 F m-1

is the permittivity of free space App is the overlapping area (m2) and d is the distance

between the two plates (m)

15

16

Figure 3-1 An inter-digitated comb drive where changes in capacitance are generated by either changes in gap distance x0 or in the overlap area y0timest

Figure 3-1

For a set of inter-digitated comb fingers or comb drive shown in the total

capacitance Ctotal is expressed as [28]

0 0

0 0

( ) ( )( ) (Total cy y y yC N t

)x x x xε

+ += + + minus

(8)

where N is the number of comb fingers on the movable comb y0 is the initial comb finger

overlap length (m) tc is the out of plane comb finger thickness (m) and x0 is the inital

finger gap (m) In this work the comb drive is implemented in three arrangements

transverse lateral and differential tri-plate In a transverse comb drive changes in

capacitance are generated by transverse movements along the x-axis resulting in changes

to the gap x0 between the movable and fixed comb fingers The nonlinear relation

between total capacitance CTotalT and displacement is expressed as

17

00 0

1 1( ) (Total TC N ty

)x x x xε

= + minus +

(9)

Although the drive sensitivity part is not constant an average sensitivity can be

assumed for small displacements [38]

Total TC xpart

In a lateral comb drive changes in capacitance are generated by lateral movements

along the y-axis resulting in changes to the overlapping area y0timestc The total capacitance

CTotalL between all comb fingers is expressed as

00

2 ( )Total LN tC yx

yε= +

(10)

The sensitivity of the lateral comb drive SL is expressed as

0

2Total LL

C N tSy x

εpart= =

part

(11)

The lateral comb drive was first introduced by W C Tang T H Nguyen and R T

Howe [33] in an electrostatic resonator and has since been used numerous times in both

drive and sense configurations It has the advantage of a fully linear response and

negligible instability but offers low sensitivity The sensitivity for the lateral drive L1

implemented in this work is 0038 pF microm-1

A differential tri-plate drive shown in Figure 3-2 which is a variation of the

transverse drive is implemented to improve sensitivity and reduce noise [39]

18

Figure 3-2 An inter-digitated comb drive in a differential tri-plate arrangement where x01 x02 and x03 are the gap distances and x01= x02ltlt x03 to improve sensitivity

In the tri-plate drive the at rest gap distances between the two capacitors formed by any

movable finger is not equivalent For a differential arrangement insuring that x01=

x02ltltx03 the linearity for small displacements (less than 1 microm) is enhanced over the case

of a single ended arrangement where x01=x02=x03=x0 which is equivalent to the

transverse drive This is shown graphically in Figure 3-3 where x01= x02=x03=2 microm in

the lower curve and x01= x02= 2 microm and x03= 15 microm in the upper curve A substantial

improvement in linearity is attained with the tri-plate drive

19

0

005

01

015

02

025

03

035

0 02 04 06 08 1

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-3 Comparison of the effect of gap distances on linearity for two case lsquorsquo represents case 1 where the fingers in are implemented in a differential tri-plate arrangement and x01= x02= 2 microm and x03= 15 microm lsquorsquo represents the case 2 where the fingers are implemented in a single ended arrangement and x01= x02=x03=2 microm

The tri-plate drive is also implemented in a differential arrangement where the output

is the change in capacitance rather than the total capacitance resulting from an external

excitation This reduces errors cause by off-axis excitation and common mode noise The

output of the differential tri-plate drive is expressed as [39]

Total Top Total BottomC C C∆ = minus where

003 01

1 1( ) (Total TopC N ty

)x x x xε

= +

minus +

(12)

20

002 03

1 1( ) (Total BottomC N ty

)x x x xε

= +

minus +

A graphic comparison of the capacitancedisplacement relation for sensors T1 and T2 is

shown in Figure 3-4 For these gap closing drives it is assumed that for displacements

less that 1 microm the relation is linear indicated by the dashed line above each curve The

sensitivity within this region is approximately 0046 pF microm-1 and 0253 pF microm-1 for T1

and T2 respectively The order of magnitude improvement is attributed to the differential

tri-plate arrangement and thicker comb finger

0

02

04

06

08

1

12

14

16

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-4 lsquorsquo represents the output of the differential tri-plate drive while lsquorsquo represents the output of the transverse drive The dashed line represents the linear approximation for displacements less that 1 microm For the purpose of comparison the output of the transverse drive was equalized to zero at zero displacement

21

In this work the transverse drive is first implemented in sensor T1 using the

PolyMUMPS fabrication process Then in an effort to linearize the sensor response the

lateral drive is used in L1 using SoiMUMPS The reduction in drive sensitivity due to the

lateral drive is further reduced by increasing the gap distance to 3 microm which is

recommended by the foundry to ensure a successful device These reductions are

balanced by increasing the number of comb fingers N and selecting a substrate thickness

that is 125 times large than T1 The third device T2 is fabricated using SoiMUMPS and

uses the tri-plate differential drive This design capitalizes on the increase in sensitivity of

a gap closing drive and larger substrate thickness attaining an order of magnitude

increase in sensitivity over T1 and L1 The specifications for each drive along with the

successive improvement in drive sensitivity are outlined in Table 3-1

Table 3-1 Specifications for each capacitive comb drive implemented in this work T1 L1 and T2

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Comb Finger Thickness tc (microm) 2 25 25 Gap Distance x0 (microm) 2 3 2 Number of Fingers N 200 256 30 Overlap length y0 (microm) 45 10 70 Drive Sensitivity (pF microm-1) 0046 0038 0273

32 MECHANICAL ANALYSIS

321 MICRO-FLEXURE SELECTION

The sensor structure consists of a proof mass connected to a compliant micro-flexure

that itself is anchored to the sensor substrate The proof mass displaces in response to an

22

applied acceleration and this displacement is sensed using the capacitive comb drive

attached to the proof mass

Proof mass

Folded beam micro-flexure

Anchor

Figure 3-5 Sensor Structure

This section focuses on micro-flexure selection as it should exhibit a large linear

displacement range and have a high lateral stiffness ratio to reduce cross-axis sensitivity

A number of micro-flexures have been used in MEMS research these include the

clamped-clamped crab-leg and folded beam shown in Figure 3-6-through 3-7

The clamped-clamped flexure exhibits a high stiffness ratio however for large lateral

displacements extensional axial forces buildup within the beams resulting in a non-linear

force-displacement relation [40] Within the linear regime the stiffness of the clamped-

clamped flexure is expressed as [40]

3

3b

yEt wk

L=

(13)

23

where E is the Youngrsquos Modulus of the beam material (N m-2) tb is the beam out of

plane thickness (m) w is the beam width (m) and L is the beam length (m) The stiffness

ratio is expressed as [40]

2x

y

k Lk w

=

(14)

If L is two orders of magnitude greater than w this results in a ratio of approximately

10000 The linear displacement range is determined using Finite Element Analysis (FEA)

of a clamped-clamped flexure with L=387 microm w=4 microm and tb=2 microm The result of this

simulation is plotted in and compared to a structure with constant stiffness ky

= 035 N m-1 predicted by Eq (13) The linear relation is valid for small deflections that

are less than 2 microm In addition this micro-flexure is susceptible to buckling as a result of

a compressive residual stress that may arise due to the fabrication process [41]

Figure 3-9

In the crab leg design the thigh section is meant to relieve the extensional axial

stress postponing the occurrence of non-linearity until greater lateral displacement [42]

The stiffness of the crab leg flexure is expressed as [43]

3

44

b s s ty

s s t

Et w L LkL L L

κκ

+= +

(15)

where κ=IsIt=(wswt)3 is a ratio of the second moment of inertia of the shin (s) and thigh

(t) ws and wt represent the width and Ls and Lt represent the length of the shin and thigh

regions respectively The stiffness ratio is expressed as [43]

34

4t s s tx

y s t s

w L L Lkk w L L L

αα t

+= +

(16)

The linear displacement range of a crab leg flexure with Ls=285 microm Lt=30 microm ws=wt=3

microm and tb=2 microm is determined using FEA The results are plotted in Figure 3-9 and

compared to a structure of constant stiffness ky=035 N m-1 predicted by Eq (15) The

24

non-linearity is greatly reduced when compared to the clamp-clamped flexure with Eq

accurately predicting the deflection until 4 microm However the flexure is more

susceptible to off-axis deflection as there is an order of magnitude reduction in the

stiffness ratio as compared to the clamped-clamped flexure

(15)

The folded beam flexure stands out as it is designed to exhibit a large linear

displacement range and to relieve residual stress due to the manufacturing process by

allowing the beams to expand and contract along the axial direction [40] It was first used

in an electrostatic resonator and has subsequently been used for electrostatic actuation in

a magnetic hard drives and micro-positioning [3344] If the truss section is rigid

ItrussgtgtIbeam then the stiffness of the flexure is expressed as [4043]

3

32b

yEt wk

L=

(17)

The stiffness ratio is expressed as [40]

2x

y

k Lk w

=

(18)

Figure 3-6 Clamped-clamped flexure

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

2

may be caused by outer bearing race defects fOB inner bearing race defects fIB ball

defects fB and train defects fT which all occur at specific frequencies [12]

cos( )12

dBOB r

p

bNf fd

β = minus

cos( )12

dBIB r

p

bNf fd

β = +

2cos( )1

2p r d

Bd p

d f bfb d

β = minus

cos( )12

drT

p

bffd

β = minus

(1)

where NB is the number of balls fr is the rotation frequency bd is the ball diameter β is the

contact angle between the ball and races and dp is the ball pitch diameter By monitoring

the real time vibration signature of the mechanical system anomalies can be quickly

identified and fixed

12 MICROELECTROMECHANICAL SENSORS

Microelectromechanical systems (MEMS) inertial sensors provide a small footprint

with sensitivities that are either comparable or exceed any macro sensor along with the

capability of mass production and low unit cost These sensors utilize compliant micro-

flexures attached to a proof mass that displaces in response to an environmental

acceleration Many transduction mechanisms have been developed that convert the

displacement into a measurable electric signal and include thermal piezoresistive

piezoelectric optical and capacitive methods Most MEMS sensors are silicon based and

are fabricated using surface or bulk micromachining Surface micromachining creates free

standing movable structures on top of a substrate using a combination of sacrificial layers

3

and structural layers which are commonly polysilicon [3] In bulk micromachining the

mechanical structures are defined using a removal process where bulk material typically

silicon is etched away MEMS sensors have been used for vibration and shock monitoring

on industrial systems and robotics guidance and navigation in global positioning systems

(GPS) seismometry in earthquake prediction image stabilization in digital cameras and

automobile safety and stability [4]

13 AUTOMOTIVE SENSORS

Sensors cover every major aspect of a modern automobile power-train sensors

monitor fuel combustion and emissions chassis sensors monitor road traction and tire

condition and body sensors facilitate air-bag deployment and vehicle proximity for radar

guided cruise control [5] Pressure sensors which typically consist of a piezoresistive

strain sensing element attached to a silicon diaphragm that deflects when exposed to an

applied pressure are one of the first micro-machined sensors used in an automobile

Implemented as manifold absolute pressure (MAP) sensors they allow precise control of

the air fuel ratio which allows the catalytic converter to efficiently reduce tailpipe

emissions [6] Variable reluctance sensors based on electromagnetics are used for

automobile traction control and produce a voltage output that is dependant on the

magnetic flux variations between a rotating component and the sensor bias magnet [57]

MEMS based linear accelerometers are utilized for airbag deployment upon impact and

provide a lower cost and smaller package over the older ball-in-tube accelerometers [6]

4

14 OPERATING ENVIRONMENT

Through proper packaging MEMS sensors can withstand the harsh automotive

environment where they face cyclic temperatures (-40degC to 150degC) mechanical shock or

vibration (10g-500g) and exposure to various fluids (oil brake fluid or engine coolant)

and airborne particles [8] Temperature changes induce thermal stresses causing

undesirable bending of the silicon microstructures that changes their dynamic response

Mechanical shock which could occur during a high impact collision or vibration from

uneven road conditions cause the delicate microstructures to fracture at points of stress

concentration These effects could be reduced by efficient design Use of a folded beam

micro-flexure that allows expansion and contraction or use of fillets at corners which

reduce the chance of fracture Bonding pads wire bonding and onboard circuit

components can be protected from air borne particles which could cause shorting using

parylene or gels Alternatively the entire package can be sealed hermetically which also

protects the MEMS device [8] In some cases such as fluid pressure sensors intimate

contact between the MEMS device and potentially corrosive fluid is necessary A recent

approach used a nanometer thin coating of silicon carbide (SiC) over a foundry fabricated

MEMS sensor using low temperature low pressure chemical vapor deposition (LPCVD)

[9] The SiC known for its chemical inertness did not compromise the mechanical

characteristics of the original design Therefore MEMS provides a strong platform for an

automotive sensor

5

15 OBJECTIVES AND CONTRIBUTION

The objective of this work is the development of a low frequency MEMS vibration

sensor The sensor must be low cost and mass producible which can be achieved using

standard foundry fabrication processes such as the Polysilicon Multi-User MEMS

Processes (PolyMUMPS) or the Silicon-on-Insulator Multi-User MEMS Processes

(SoiMUMPS) In addition the number of additional processing steps which are required

if integrated CMOS circuitry is used or if vacuum packaging is required must be

minimized This can be achieved using an off-chip circuit made with widely available

discrete components and sensor operation in air

In addition this work presents a MEMS sensor that is fabricated using standard

foundry processes such as PolyMUMPS and SoiMUMPS compared to many MEMS

sensors in literature that are fabricated using customized processes

16 RESEARCH OUTLINE

This work demonstrates three iterations of a low frequency capacitive vibration

sensor fabricated using the standard foundry processes Each sensor consists of a

suspended proof mass that displaces in response to an external vibration The proof mass

is attached to a folded beam micro-flexure compliant in one direction and its lateral

movement is sensed using a capacitive comb drive The first sensor T1 is fabricated in

PolyMUMPS and uses a transverse capacitive comb drive The second L1 fabricated in

SoiMUMPS uses a fully linear lateral capacitive drive and achieves an increase in

sensitivity from T1 Finally the third sensor T2 is fabricated in SoiMUMPS and uses a

differential tri-plate capacitive drive This sensor takes advantage of the increase in

6

sensitivity gained from a transverse drive but rejects off-axis displacement errors and

common mode using a differential tri-plate arrangement In addition an off-chip

capacitance to voltage readout circuit is fabricated and tested on sensor L1

Chapter 2 begins with an overview on the various methods of fault detection

followed by a review of MEMS vibration sensors each with a different transduction

mechanism Chapter 3 outlines the sensor structure focusing on the mechanical design

and sensing principal Chapter 4 focuses on the readout circuitry layout and simulation

results Chapter 5 outlines the various source of noise including mechanical noise arising

from the sensor and electric noise from the readout circuit Chapter 6 overviews the

foundry fabrication processes used including PolyMUMPs and SoiMUMPS The

experimental results of two sensors T1 and L1 which were fabricated and tested within

the thesis time frame are presented in Chapter 7 Chapter 8 concludes the thesis with a

brief summary and outline for future work

2 LITERATURE REVIEW

21 EXISTING METHODS FOR FAULT DETECTION

As this thesis develops a sensor for fault detection using vibration monitoring this

section overviews three existing methods for fault detection in mechanical systems First

motor current monitoring is discussed as it has the ability to measure significant bearing

faults in motor systems Next temperature sensors specifically thin film thermocouples

(TFTC) and MEMS temperature sensors are discussed These sensors are mounted in the

chamber of a lubricating fluid or in close proximity to a moving component and monitor

temperature fluctuations caused by increased wear Finally acoustic emission (AE)

sensors are discussed which have the ability to detect ultrasonic emissions generated by

plastic deformation in the crystalline lattice of a mechanical component

211 MOTOR CURRENT MONITORING

Monitoring of stator current is used as a method of determining bearing related faults

that produce radial motion between rotor and stator This motion varies the air gap flux

density producing stator currents at predictable frequencies fbng that represent the

bearing faults [10]

bng e vf f m f= plusmn sdot (2)

where fe is the electrical supply frequency m=123hellip and fv is one of the characteristics

frequencies listed in the equations of (1)The bearing damage must be significant to

produce noticeable current fluctuations and it may be difficult to distinguish from current

fluctuations due to voltage imbalances machine saturation or statorarmature faults [2]

7

8

212 TEMPERATURE SENSORS

Temperature sensors are largely used in high temperature high load and high speed

environments typically turbine engines and industrial tools Thermocouple sensors

consist of two dissimilar metals joined together at two junctions and when exposed to a

temperature difference generate a thermoelectric voltage dependant only on the material

properties and junction temperature difference [11] In turbine engines these sensors can

be placed on bearing casings or in the lubricant to monitor temperature fluctuations due

to increased friction The thermocouples are fabricated using bulk materials or micro-

machined thin films (TFTC) with the latter providing a faster response small package

and substantially lower production costs [12] The TFTC can be embedded for in situ

monitoring of processes in hostile environments One study successfully embedded

TFTC into electroplated nickel and attained device characteristics on par with surface

mounted devices [13] More recently MEMS temperature sensors have been developed

that consist of micro-machined semiconductor material that undergoes structural

deformation with temperature changes A common implementation is a bent beam

structure whose temperature induced deflection δ is expressed as [14]

( ) 3

2 2

sin12 cos sinz

EA T LEI AEL

α βδ 2β β

∆=

+

(3)

Temperature sensors are well suited for localized fault monitoring and detection but can

only characterize a mechanical system consisting of multiple components through the use

of multiple sensors

9

213 ACOUSTIC EMISSION SENSORS

Acoustic emission (AE) sensors have been used to characterize wear in machine

tools and monitor bearing and gear problems in centrifugal pumps [1516] First

developed as a Non-Destructive Testing (NDT) technique to detect cracks in civil

structures these sensors detect acoustic emissions generated by the release of vibration

waves in a crystalline lattice due to plastic deformation or crack growth [16]

Measurements are made using piezoelectric transducers with high natural frequencies

100 kHz to 1 MHz to capture the ultrasonic AE emissions An AE sensor is useful as it

has the ability to detect subsurface cracks in gear teeth or bearings before appearing on

the surface causing further damage [17] More recently MEMS acoustic sensors have

been developed and one design includes multiple transducers on a single substrate

which each detect acoustic emission energy at different frequencies [18] This helps

distinguish spurious acoustic emissions arising from impact and friction from those

arising from plastic deformation When compared to typical piezoelectric sensors the

MEMS devices have lower sensitivities and fail to detect some acoustic emissions

[1819] In addition the acoustic emission signal suffers severe attenuation as is crosses

various interfaces such as a gearbox or bearing casings In one experiment consisting of

a pinion gear and an associated bearing a 44dB attenuation was seen between an AE

sensor placed directly on the pinion to one placed on the bearing casing and in some

cases intermediate loss of the signal was observed [20]

10

22 MEMS VIBRATION SENSORS

This thesis focuses on the development of a MEMS vibration sensor for fault

detection as the vibration signature of a mechanical system has the potential to give an

overview of the entire system Many types of MEMS vibration sensors exist with

different operating principals based on resonance the piezoelectric effect and

displacement variation

221 RESONANT SENSORS

Resonant sensors are based on the resonant frequency shift of a fixed-fixed micro-

beam that is excited into resonance using electrostatic thermal or piezoelectric methods

Under an applied axial load there is a shift in the resonant frequency which is expressed

as [21]

2 2

2

1 12 1

nn n

EI Fl2

fl A Eα γ

π ρ= +

I

(4)

where l is the beam length E is Youngrsquos Modulus I is the moment of inertia ρ is the

density A is the cross section area F is the applied axial force αn and γn are mode

dependant coefficients and are 4730 and 0295 for mode 1 in the case discussed The

axial force F is generated by a proof mass attached to one beam end which moves in

response to an external acceleration The performance of resonant sensors are dependant

on the quality factor Q which is the ratio of total energy stored in the system to the

energy lost per cycle and determines the sharpness and amplitude of the resonance peak

A high Q resonator develops a peak that is easily distinguishable in a phase locked loop

control circuit and allows for improved performance and resolution [3] A resonator Q

value is lowered by energy losses to the surrounding fluid attached supports and

11

material Losses to the fluid surrounding the MEMS structure is the dominant loss

mechanism for MEMS sensors operated in air and can be greatly reduced by vacuum

packaging which can be accomplished using glass-silicon anodic bonding [22] In

addition to increasing fabrication cost the long term stability of the vacuum degrades

with ultimate device failure stemming from leakage through micro cracks and defects

and out-gassing [22] Support losses arise from the restoring forces they generate and

can be reduce by balanced designs such as a double ended tuning fork (DETF) or

operation in higher modes [3] In addition the sensitivity of a resonant sensor can be

increased by the addition of micro-levers to increase the axial force In one such design a

two stage micro-lever increased the sensitivity by an order of magnitude over a more

conventional single lever arm [2324]

222 PIEZOELECTRIC SENSORS

Piezoelectric materials produce a charge in response to an applied force This is an

intrinsic effect in materials such as Zinc Oxide (ZnO) whereas materials such as Barium

Titanate or Lead Zirconate Titanate (PZT) need to be poled by placing the material in a

strong electric field at an elevated temperature [3] The piezoelectric effect is expressed

mathematical by [25]

i ij j ik kD d T Eε= + (5)

where D represents the electrical displacement (Cm2) d is the charge coefficient (CN)

T is the applied stress matrix (Nm2) ε is the permittivity matrix (Fm) and E is the

electric field matrix (Vm) The subscripts i=123 j=123456 and k=123 represent

the direction to which the physical property is related For micro-sensors poled in the out

of plane direction only the D3 term is relevant and if fabricated using ZnO or PZT only

12

the charge coefficients d31 d32 and d33 are relevant [25] The piezoelectric material is

commonly deposited as a thin film on the surface of compliant structures in bulk micro-

machined sensors or incorporated as a layer in surface micro-machining on a similar

structure [26] However the inclusion of multiple materials introduces large temperature

variations causing out of plane bending especially in surface micro-machined devices

[2627] Poled piezoelectric materials could be depolarized which serves to reduce the

piezoelectric affect in the material This can occur if the sensor is exposed to a strong

electric field of the opposite polarity high temperatures in excess of the Curie point or

high mechanical stress Also the piezoelectric effect reduces over time an effect that can

be reduced with the addition of composite elements [28] Foundry processes such as

PolyMUMPS and SoiMUMPS do not incorporate piezoelectric layers thus sensors using

this method of transduction must use a customized process The piezoelectric sensor

outlined in [26] uses polysilicon surface micro-machining similar to PolyMUMPS

however it incorporates a ZnO layer that is not available in the foundry process

223 DISPLACEMENT VARIATION SENSORS

Displacement variation sensors consist of a proof mass connected to a compliant

micro-flexure The movement of the proof mass is sensed using optical electron

tunneling or more commonly capacitive methods A recent optical accelerometer

utilized a novel multilayer nano-grating for in-plane displacement sensing of the proof

mass [29] For small space variations between the nano-gratings a large change in the

optical reflection amplitude of the grating was observed [30] Even though optical

sensors offer resolutions approaching the Brownian noise limit they are not available in

small packages that are easily placed in space constricted areas for a low cost [29]

13

The electron tunneling effect is observed when the proximity between two electrodes

one flexible and the other fixed is on the order of 10Ǻ [31] If the two electrodes are

biased with voltage VB a tunneling current It flows between the gap [31]

exp( )t B I tgI V xαinfin sdot minus Φ (6)

where α1 is a constant and equal to 1025 Ǻ-1 eV-frac12 Ф is the effective height of the

tunneling barrier and xtg is the tunneling gap When the flexible electrode moves in

response to acceleration a feedback controller maintains the original distance while

determining the applied acceleration Electron tunneling sensors are able to sense

accelerations in the microg-ng range due to the exponential relation between tunneling current

and gap However these highly sensitive devices are susceptible to thermo-mechanical

noise which could be reduced by operation at low pressure [32] It is also important for

the two electrodes to be coated with metal which is difficult to achieve with foundry

fabrication

Capacitive sensors are based on the parallel plate capacitor and are implemented

using inter-digitated comb fingers in an in-plane lateral or transverse drive or in an out of

plane parallel plate drive [33] These sensors have the advantage of high sensitivity low

power consumption small package low temperature dependence and are easily

integrated with existing foundry fabrication processes [34] The lateral drive offers a fully

linear response but low sensitivity whereas the transverse and parallel plate drives

achieve high sensitivity if gap distances are of a few microns and large capacitance areas

are utilized However both drives have a non-linear response and are prone to pull-in

Also if these sensors are fabricated using a surface micromachining process such as

PolyMUMPS where the ground plane is a few microns above the structural layer they

14

are prone to electrical failure due to shorting This was observed in an experiment where

similar sensors were placed in a vibrating environment with a peak of 120g and shorting

between the structural layer and ground plane was found to be the predominate mode of

failure [35] However with certain fabrication processes such as SoiMUMPS this issue

can be neglected as the substrate material below any moving component is removed

Most capacitance sensors are fabricated using an integrated CMOS (Complementary

Metal Oxide Semiconductor)-MEMS technique where the MEMS sensor is fabricated

before during or after the CMOS circuit fabrication [36] Although these sensors offer

high-sensitivity and small parasitic capacitance they suffer from in-plane and out of

plane curling of the beam which reduces the capacitance between adjacent comb fingers

[37] In addition multiple processing steps are required for device fabrication

More than the other transduction methods MEMS capacitive displacement sensors are

attractive for practical implementation in an automobile and therefore is the selected

method for the vibration sensor presented in this work Their low temperature dependence

is an ideal characteristic for an automobile environment that faces varying temperature If

implemented with off-chip readout circuitry they require a standard foundry fabrication

process which provides a strong platform for mass production and low unit cost

3 SENSOR STRUCTURE

This chapter presents the physical structure of the three MEMS vibration sensors

which includes the displacement sensing capacitive comb drive and the micro-

flexureproof mass system First the capacitive comb drive and the three variations

implemented the transverse drive lateral drive and differential tri-plate drive are

discussed Next the mechanical structure is presented and includes the micro-flexure

selection quality factor and equations of motion Finally instability that places a limit

on the operating range of the sensor as a result of a gap closing capacitive comb drive is

discussed

31 SENSING

Capacitive sensors are based on the parallel plate capacitor and are commonly

implemented using inter-digitated comb fingers A capacitance can be realized if a voltage

is placed between two closely spaced plates If the fringing electric field is neglected the

capacitance of the parallel plate capacitor Cpp (F) is expressed as [28]

air pppp

k AC

dε

= (7)

where kair=1 is the dielectric constant of air at atmospheric pressure ε=88542times10-12 F m-1

is the permittivity of free space App is the overlapping area (m2) and d is the distance

between the two plates (m)

15

16

Figure 3-1 An inter-digitated comb drive where changes in capacitance are generated by either changes in gap distance x0 or in the overlap area y0timest

Figure 3-1

For a set of inter-digitated comb fingers or comb drive shown in the total

capacitance Ctotal is expressed as [28]

0 0

0 0

( ) ( )( ) (Total cy y y yC N t

)x x x xε

+ += + + minus

(8)

where N is the number of comb fingers on the movable comb y0 is the initial comb finger

overlap length (m) tc is the out of plane comb finger thickness (m) and x0 is the inital

finger gap (m) In this work the comb drive is implemented in three arrangements

transverse lateral and differential tri-plate In a transverse comb drive changes in

capacitance are generated by transverse movements along the x-axis resulting in changes

to the gap x0 between the movable and fixed comb fingers The nonlinear relation

between total capacitance CTotalT and displacement is expressed as

17

00 0

1 1( ) (Total TC N ty

)x x x xε

= + minus +

(9)

Although the drive sensitivity part is not constant an average sensitivity can be

assumed for small displacements [38]

Total TC xpart

In a lateral comb drive changes in capacitance are generated by lateral movements

along the y-axis resulting in changes to the overlapping area y0timestc The total capacitance

CTotalL between all comb fingers is expressed as

00

2 ( )Total LN tC yx

yε= +

(10)

The sensitivity of the lateral comb drive SL is expressed as

0

2Total LL

C N tSy x

εpart= =

part

(11)

The lateral comb drive was first introduced by W C Tang T H Nguyen and R T

Howe [33] in an electrostatic resonator and has since been used numerous times in both

drive and sense configurations It has the advantage of a fully linear response and

negligible instability but offers low sensitivity The sensitivity for the lateral drive L1

implemented in this work is 0038 pF microm-1

A differential tri-plate drive shown in Figure 3-2 which is a variation of the

transverse drive is implemented to improve sensitivity and reduce noise [39]

18

Figure 3-2 An inter-digitated comb drive in a differential tri-plate arrangement where x01 x02 and x03 are the gap distances and x01= x02ltlt x03 to improve sensitivity

In the tri-plate drive the at rest gap distances between the two capacitors formed by any

movable finger is not equivalent For a differential arrangement insuring that x01=

x02ltltx03 the linearity for small displacements (less than 1 microm) is enhanced over the case

of a single ended arrangement where x01=x02=x03=x0 which is equivalent to the

transverse drive This is shown graphically in Figure 3-3 where x01= x02=x03=2 microm in

the lower curve and x01= x02= 2 microm and x03= 15 microm in the upper curve A substantial

improvement in linearity is attained with the tri-plate drive

19

0

005

01

015

02

025

03

035

0 02 04 06 08 1

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-3 Comparison of the effect of gap distances on linearity for two case lsquorsquo represents case 1 where the fingers in are implemented in a differential tri-plate arrangement and x01= x02= 2 microm and x03= 15 microm lsquorsquo represents the case 2 where the fingers are implemented in a single ended arrangement and x01= x02=x03=2 microm

The tri-plate drive is also implemented in a differential arrangement where the output

is the change in capacitance rather than the total capacitance resulting from an external

excitation This reduces errors cause by off-axis excitation and common mode noise The

output of the differential tri-plate drive is expressed as [39]

Total Top Total BottomC C C∆ = minus where

003 01

1 1( ) (Total TopC N ty

)x x x xε

= +

minus +

(12)

20

002 03

1 1( ) (Total BottomC N ty

)x x x xε

= +

minus +

A graphic comparison of the capacitancedisplacement relation for sensors T1 and T2 is

shown in Figure 3-4 For these gap closing drives it is assumed that for displacements

less that 1 microm the relation is linear indicated by the dashed line above each curve The

sensitivity within this region is approximately 0046 pF microm-1 and 0253 pF microm-1 for T1

and T2 respectively The order of magnitude improvement is attributed to the differential

tri-plate arrangement and thicker comb finger

0

02

04

06

08

1

12

14

16

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-4 lsquorsquo represents the output of the differential tri-plate drive while lsquorsquo represents the output of the transverse drive The dashed line represents the linear approximation for displacements less that 1 microm For the purpose of comparison the output of the transverse drive was equalized to zero at zero displacement

21

In this work the transverse drive is first implemented in sensor T1 using the

PolyMUMPS fabrication process Then in an effort to linearize the sensor response the

lateral drive is used in L1 using SoiMUMPS The reduction in drive sensitivity due to the

lateral drive is further reduced by increasing the gap distance to 3 microm which is

recommended by the foundry to ensure a successful device These reductions are

balanced by increasing the number of comb fingers N and selecting a substrate thickness

that is 125 times large than T1 The third device T2 is fabricated using SoiMUMPS and

uses the tri-plate differential drive This design capitalizes on the increase in sensitivity of

a gap closing drive and larger substrate thickness attaining an order of magnitude

increase in sensitivity over T1 and L1 The specifications for each drive along with the

successive improvement in drive sensitivity are outlined in Table 3-1

Table 3-1 Specifications for each capacitive comb drive implemented in this work T1 L1 and T2

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Comb Finger Thickness tc (microm) 2 25 25 Gap Distance x0 (microm) 2 3 2 Number of Fingers N 200 256 30 Overlap length y0 (microm) 45 10 70 Drive Sensitivity (pF microm-1) 0046 0038 0273

32 MECHANICAL ANALYSIS

321 MICRO-FLEXURE SELECTION

The sensor structure consists of a proof mass connected to a compliant micro-flexure

that itself is anchored to the sensor substrate The proof mass displaces in response to an

22

applied acceleration and this displacement is sensed using the capacitive comb drive

attached to the proof mass

Proof mass

Folded beam micro-flexure

Anchor

Figure 3-5 Sensor Structure

This section focuses on micro-flexure selection as it should exhibit a large linear

displacement range and have a high lateral stiffness ratio to reduce cross-axis sensitivity

A number of micro-flexures have been used in MEMS research these include the

clamped-clamped crab-leg and folded beam shown in Figure 3-6-through 3-7

The clamped-clamped flexure exhibits a high stiffness ratio however for large lateral

displacements extensional axial forces buildup within the beams resulting in a non-linear

force-displacement relation [40] Within the linear regime the stiffness of the clamped-

clamped flexure is expressed as [40]

3

3b

yEt wk

L=

(13)

23

where E is the Youngrsquos Modulus of the beam material (N m-2) tb is the beam out of

plane thickness (m) w is the beam width (m) and L is the beam length (m) The stiffness

ratio is expressed as [40]

2x

y

k Lk w

=

(14)

If L is two orders of magnitude greater than w this results in a ratio of approximately

10000 The linear displacement range is determined using Finite Element Analysis (FEA)

of a clamped-clamped flexure with L=387 microm w=4 microm and tb=2 microm The result of this

simulation is plotted in and compared to a structure with constant stiffness ky

= 035 N m-1 predicted by Eq (13) The linear relation is valid for small deflections that

are less than 2 microm In addition this micro-flexure is susceptible to buckling as a result of

a compressive residual stress that may arise due to the fabrication process [41]

Figure 3-9

In the crab leg design the thigh section is meant to relieve the extensional axial

stress postponing the occurrence of non-linearity until greater lateral displacement [42]

The stiffness of the crab leg flexure is expressed as [43]

3

44

b s s ty

s s t

Et w L LkL L L

κκ

+= +

(15)

where κ=IsIt=(wswt)3 is a ratio of the second moment of inertia of the shin (s) and thigh

(t) ws and wt represent the width and Ls and Lt represent the length of the shin and thigh

regions respectively The stiffness ratio is expressed as [43]

34

4t s s tx

y s t s

w L L Lkk w L L L

αα t

+= +

(16)

The linear displacement range of a crab leg flexure with Ls=285 microm Lt=30 microm ws=wt=3

microm and tb=2 microm is determined using FEA The results are plotted in Figure 3-9 and

compared to a structure of constant stiffness ky=035 N m-1 predicted by Eq (15) The

24

non-linearity is greatly reduced when compared to the clamp-clamped flexure with Eq

accurately predicting the deflection until 4 microm However the flexure is more

susceptible to off-axis deflection as there is an order of magnitude reduction in the

stiffness ratio as compared to the clamped-clamped flexure

(15)

The folded beam flexure stands out as it is designed to exhibit a large linear

displacement range and to relieve residual stress due to the manufacturing process by

allowing the beams to expand and contract along the axial direction [40] It was first used

in an electrostatic resonator and has subsequently been used for electrostatic actuation in

a magnetic hard drives and micro-positioning [3344] If the truss section is rigid

ItrussgtgtIbeam then the stiffness of the flexure is expressed as [4043]

3

32b

yEt wk

L=

(17)

The stiffness ratio is expressed as [40]

2x

y

k Lk w

=

(18)

Figure 3-6 Clamped-clamped flexure

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

3

and structural layers which are commonly polysilicon [3] In bulk micromachining the

mechanical structures are defined using a removal process where bulk material typically

silicon is etched away MEMS sensors have been used for vibration and shock monitoring

on industrial systems and robotics guidance and navigation in global positioning systems

(GPS) seismometry in earthquake prediction image stabilization in digital cameras and

automobile safety and stability [4]

13 AUTOMOTIVE SENSORS

Sensors cover every major aspect of a modern automobile power-train sensors

monitor fuel combustion and emissions chassis sensors monitor road traction and tire

condition and body sensors facilitate air-bag deployment and vehicle proximity for radar

guided cruise control [5] Pressure sensors which typically consist of a piezoresistive

strain sensing element attached to a silicon diaphragm that deflects when exposed to an

applied pressure are one of the first micro-machined sensors used in an automobile

Implemented as manifold absolute pressure (MAP) sensors they allow precise control of

the air fuel ratio which allows the catalytic converter to efficiently reduce tailpipe

emissions [6] Variable reluctance sensors based on electromagnetics are used for

automobile traction control and produce a voltage output that is dependant on the

magnetic flux variations between a rotating component and the sensor bias magnet [57]

MEMS based linear accelerometers are utilized for airbag deployment upon impact and

provide a lower cost and smaller package over the older ball-in-tube accelerometers [6]

4

14 OPERATING ENVIRONMENT

Through proper packaging MEMS sensors can withstand the harsh automotive

environment where they face cyclic temperatures (-40degC to 150degC) mechanical shock or

vibration (10g-500g) and exposure to various fluids (oil brake fluid or engine coolant)

and airborne particles [8] Temperature changes induce thermal stresses causing

undesirable bending of the silicon microstructures that changes their dynamic response

Mechanical shock which could occur during a high impact collision or vibration from

uneven road conditions cause the delicate microstructures to fracture at points of stress

concentration These effects could be reduced by efficient design Use of a folded beam

micro-flexure that allows expansion and contraction or use of fillets at corners which

reduce the chance of fracture Bonding pads wire bonding and onboard circuit

components can be protected from air borne particles which could cause shorting using

parylene or gels Alternatively the entire package can be sealed hermetically which also

protects the MEMS device [8] In some cases such as fluid pressure sensors intimate

contact between the MEMS device and potentially corrosive fluid is necessary A recent

approach used a nanometer thin coating of silicon carbide (SiC) over a foundry fabricated

MEMS sensor using low temperature low pressure chemical vapor deposition (LPCVD)

[9] The SiC known for its chemical inertness did not compromise the mechanical

characteristics of the original design Therefore MEMS provides a strong platform for an

automotive sensor

5

15 OBJECTIVES AND CONTRIBUTION

The objective of this work is the development of a low frequency MEMS vibration

sensor The sensor must be low cost and mass producible which can be achieved using

standard foundry fabrication processes such as the Polysilicon Multi-User MEMS

Processes (PolyMUMPS) or the Silicon-on-Insulator Multi-User MEMS Processes

(SoiMUMPS) In addition the number of additional processing steps which are required

if integrated CMOS circuitry is used or if vacuum packaging is required must be

minimized This can be achieved using an off-chip circuit made with widely available

discrete components and sensor operation in air

In addition this work presents a MEMS sensor that is fabricated using standard

foundry processes such as PolyMUMPS and SoiMUMPS compared to many MEMS

sensors in literature that are fabricated using customized processes

16 RESEARCH OUTLINE

This work demonstrates three iterations of a low frequency capacitive vibration

sensor fabricated using the standard foundry processes Each sensor consists of a

suspended proof mass that displaces in response to an external vibration The proof mass

is attached to a folded beam micro-flexure compliant in one direction and its lateral

movement is sensed using a capacitive comb drive The first sensor T1 is fabricated in

PolyMUMPS and uses a transverse capacitive comb drive The second L1 fabricated in

SoiMUMPS uses a fully linear lateral capacitive drive and achieves an increase in

sensitivity from T1 Finally the third sensor T2 is fabricated in SoiMUMPS and uses a

differential tri-plate capacitive drive This sensor takes advantage of the increase in

6

sensitivity gained from a transverse drive but rejects off-axis displacement errors and

common mode using a differential tri-plate arrangement In addition an off-chip

capacitance to voltage readout circuit is fabricated and tested on sensor L1

Chapter 2 begins with an overview on the various methods of fault detection

followed by a review of MEMS vibration sensors each with a different transduction

mechanism Chapter 3 outlines the sensor structure focusing on the mechanical design

and sensing principal Chapter 4 focuses on the readout circuitry layout and simulation

results Chapter 5 outlines the various source of noise including mechanical noise arising

from the sensor and electric noise from the readout circuit Chapter 6 overviews the

foundry fabrication processes used including PolyMUMPs and SoiMUMPS The

experimental results of two sensors T1 and L1 which were fabricated and tested within

the thesis time frame are presented in Chapter 7 Chapter 8 concludes the thesis with a

brief summary and outline for future work

2 LITERATURE REVIEW

21 EXISTING METHODS FOR FAULT DETECTION

As this thesis develops a sensor for fault detection using vibration monitoring this

section overviews three existing methods for fault detection in mechanical systems First

motor current monitoring is discussed as it has the ability to measure significant bearing

faults in motor systems Next temperature sensors specifically thin film thermocouples

(TFTC) and MEMS temperature sensors are discussed These sensors are mounted in the

chamber of a lubricating fluid or in close proximity to a moving component and monitor

temperature fluctuations caused by increased wear Finally acoustic emission (AE)

sensors are discussed which have the ability to detect ultrasonic emissions generated by

plastic deformation in the crystalline lattice of a mechanical component

211 MOTOR CURRENT MONITORING

Monitoring of stator current is used as a method of determining bearing related faults

that produce radial motion between rotor and stator This motion varies the air gap flux

density producing stator currents at predictable frequencies fbng that represent the

bearing faults [10]

bng e vf f m f= plusmn sdot (2)

where fe is the electrical supply frequency m=123hellip and fv is one of the characteristics

frequencies listed in the equations of (1)The bearing damage must be significant to

produce noticeable current fluctuations and it may be difficult to distinguish from current

fluctuations due to voltage imbalances machine saturation or statorarmature faults [2]

7

8

212 TEMPERATURE SENSORS

Temperature sensors are largely used in high temperature high load and high speed

environments typically turbine engines and industrial tools Thermocouple sensors

consist of two dissimilar metals joined together at two junctions and when exposed to a

temperature difference generate a thermoelectric voltage dependant only on the material

properties and junction temperature difference [11] In turbine engines these sensors can

be placed on bearing casings or in the lubricant to monitor temperature fluctuations due

to increased friction The thermocouples are fabricated using bulk materials or micro-

machined thin films (TFTC) with the latter providing a faster response small package

and substantially lower production costs [12] The TFTC can be embedded for in situ

monitoring of processes in hostile environments One study successfully embedded

TFTC into electroplated nickel and attained device characteristics on par with surface

mounted devices [13] More recently MEMS temperature sensors have been developed

that consist of micro-machined semiconductor material that undergoes structural

deformation with temperature changes A common implementation is a bent beam

structure whose temperature induced deflection δ is expressed as [14]

( ) 3

2 2

sin12 cos sinz

EA T LEI AEL

α βδ 2β β

∆=

+

(3)

Temperature sensors are well suited for localized fault monitoring and detection but can

only characterize a mechanical system consisting of multiple components through the use

of multiple sensors

9

213 ACOUSTIC EMISSION SENSORS

Acoustic emission (AE) sensors have been used to characterize wear in machine

tools and monitor bearing and gear problems in centrifugal pumps [1516] First

developed as a Non-Destructive Testing (NDT) technique to detect cracks in civil

structures these sensors detect acoustic emissions generated by the release of vibration

waves in a crystalline lattice due to plastic deformation or crack growth [16]

Measurements are made using piezoelectric transducers with high natural frequencies

100 kHz to 1 MHz to capture the ultrasonic AE emissions An AE sensor is useful as it

has the ability to detect subsurface cracks in gear teeth or bearings before appearing on

the surface causing further damage [17] More recently MEMS acoustic sensors have

been developed and one design includes multiple transducers on a single substrate

which each detect acoustic emission energy at different frequencies [18] This helps

distinguish spurious acoustic emissions arising from impact and friction from those

arising from plastic deformation When compared to typical piezoelectric sensors the

MEMS devices have lower sensitivities and fail to detect some acoustic emissions

[1819] In addition the acoustic emission signal suffers severe attenuation as is crosses

various interfaces such as a gearbox or bearing casings In one experiment consisting of

a pinion gear and an associated bearing a 44dB attenuation was seen between an AE

sensor placed directly on the pinion to one placed on the bearing casing and in some

cases intermediate loss of the signal was observed [20]

10

22 MEMS VIBRATION SENSORS

This thesis focuses on the development of a MEMS vibration sensor for fault

detection as the vibration signature of a mechanical system has the potential to give an

overview of the entire system Many types of MEMS vibration sensors exist with

different operating principals based on resonance the piezoelectric effect and

displacement variation

221 RESONANT SENSORS

Resonant sensors are based on the resonant frequency shift of a fixed-fixed micro-

beam that is excited into resonance using electrostatic thermal or piezoelectric methods

Under an applied axial load there is a shift in the resonant frequency which is expressed

as [21]

2 2

2

1 12 1

nn n

EI Fl2

fl A Eα γ

π ρ= +

I

(4)

where l is the beam length E is Youngrsquos Modulus I is the moment of inertia ρ is the

density A is the cross section area F is the applied axial force αn and γn are mode

dependant coefficients and are 4730 and 0295 for mode 1 in the case discussed The

axial force F is generated by a proof mass attached to one beam end which moves in

response to an external acceleration The performance of resonant sensors are dependant

on the quality factor Q which is the ratio of total energy stored in the system to the

energy lost per cycle and determines the sharpness and amplitude of the resonance peak

A high Q resonator develops a peak that is easily distinguishable in a phase locked loop

control circuit and allows for improved performance and resolution [3] A resonator Q

value is lowered by energy losses to the surrounding fluid attached supports and

11

material Losses to the fluid surrounding the MEMS structure is the dominant loss

mechanism for MEMS sensors operated in air and can be greatly reduced by vacuum

packaging which can be accomplished using glass-silicon anodic bonding [22] In

addition to increasing fabrication cost the long term stability of the vacuum degrades

with ultimate device failure stemming from leakage through micro cracks and defects

and out-gassing [22] Support losses arise from the restoring forces they generate and

can be reduce by balanced designs such as a double ended tuning fork (DETF) or

operation in higher modes [3] In addition the sensitivity of a resonant sensor can be

increased by the addition of micro-levers to increase the axial force In one such design a

two stage micro-lever increased the sensitivity by an order of magnitude over a more

conventional single lever arm [2324]

222 PIEZOELECTRIC SENSORS

Piezoelectric materials produce a charge in response to an applied force This is an

intrinsic effect in materials such as Zinc Oxide (ZnO) whereas materials such as Barium

Titanate or Lead Zirconate Titanate (PZT) need to be poled by placing the material in a

strong electric field at an elevated temperature [3] The piezoelectric effect is expressed

mathematical by [25]

i ij j ik kD d T Eε= + (5)

where D represents the electrical displacement (Cm2) d is the charge coefficient (CN)

T is the applied stress matrix (Nm2) ε is the permittivity matrix (Fm) and E is the

electric field matrix (Vm) The subscripts i=123 j=123456 and k=123 represent

the direction to which the physical property is related For micro-sensors poled in the out

of plane direction only the D3 term is relevant and if fabricated using ZnO or PZT only

12

the charge coefficients d31 d32 and d33 are relevant [25] The piezoelectric material is

commonly deposited as a thin film on the surface of compliant structures in bulk micro-

machined sensors or incorporated as a layer in surface micro-machining on a similar

structure [26] However the inclusion of multiple materials introduces large temperature

variations causing out of plane bending especially in surface micro-machined devices

[2627] Poled piezoelectric materials could be depolarized which serves to reduce the

piezoelectric affect in the material This can occur if the sensor is exposed to a strong

electric field of the opposite polarity high temperatures in excess of the Curie point or

high mechanical stress Also the piezoelectric effect reduces over time an effect that can

be reduced with the addition of composite elements [28] Foundry processes such as

PolyMUMPS and SoiMUMPS do not incorporate piezoelectric layers thus sensors using

this method of transduction must use a customized process The piezoelectric sensor

outlined in [26] uses polysilicon surface micro-machining similar to PolyMUMPS

however it incorporates a ZnO layer that is not available in the foundry process

223 DISPLACEMENT VARIATION SENSORS

Displacement variation sensors consist of a proof mass connected to a compliant

micro-flexure The movement of the proof mass is sensed using optical electron

tunneling or more commonly capacitive methods A recent optical accelerometer

utilized a novel multilayer nano-grating for in-plane displacement sensing of the proof

mass [29] For small space variations between the nano-gratings a large change in the

optical reflection amplitude of the grating was observed [30] Even though optical

sensors offer resolutions approaching the Brownian noise limit they are not available in

small packages that are easily placed in space constricted areas for a low cost [29]

13

The electron tunneling effect is observed when the proximity between two electrodes

one flexible and the other fixed is on the order of 10Ǻ [31] If the two electrodes are

biased with voltage VB a tunneling current It flows between the gap [31]

exp( )t B I tgI V xαinfin sdot minus Φ (6)

where α1 is a constant and equal to 1025 Ǻ-1 eV-frac12 Ф is the effective height of the

tunneling barrier and xtg is the tunneling gap When the flexible electrode moves in

response to acceleration a feedback controller maintains the original distance while

determining the applied acceleration Electron tunneling sensors are able to sense

accelerations in the microg-ng range due to the exponential relation between tunneling current

and gap However these highly sensitive devices are susceptible to thermo-mechanical

noise which could be reduced by operation at low pressure [32] It is also important for

the two electrodes to be coated with metal which is difficult to achieve with foundry

fabrication

Capacitive sensors are based on the parallel plate capacitor and are implemented

using inter-digitated comb fingers in an in-plane lateral or transverse drive or in an out of

plane parallel plate drive [33] These sensors have the advantage of high sensitivity low

power consumption small package low temperature dependence and are easily

integrated with existing foundry fabrication processes [34] The lateral drive offers a fully

linear response but low sensitivity whereas the transverse and parallel plate drives

achieve high sensitivity if gap distances are of a few microns and large capacitance areas

are utilized However both drives have a non-linear response and are prone to pull-in

Also if these sensors are fabricated using a surface micromachining process such as

PolyMUMPS where the ground plane is a few microns above the structural layer they

14

are prone to electrical failure due to shorting This was observed in an experiment where

similar sensors were placed in a vibrating environment with a peak of 120g and shorting

between the structural layer and ground plane was found to be the predominate mode of

failure [35] However with certain fabrication processes such as SoiMUMPS this issue

can be neglected as the substrate material below any moving component is removed

Most capacitance sensors are fabricated using an integrated CMOS (Complementary

Metal Oxide Semiconductor)-MEMS technique where the MEMS sensor is fabricated

before during or after the CMOS circuit fabrication [36] Although these sensors offer

high-sensitivity and small parasitic capacitance they suffer from in-plane and out of

plane curling of the beam which reduces the capacitance between adjacent comb fingers

[37] In addition multiple processing steps are required for device fabrication

More than the other transduction methods MEMS capacitive displacement sensors are

attractive for practical implementation in an automobile and therefore is the selected

method for the vibration sensor presented in this work Their low temperature dependence

is an ideal characteristic for an automobile environment that faces varying temperature If

implemented with off-chip readout circuitry they require a standard foundry fabrication

process which provides a strong platform for mass production and low unit cost

3 SENSOR STRUCTURE

This chapter presents the physical structure of the three MEMS vibration sensors

which includes the displacement sensing capacitive comb drive and the micro-

flexureproof mass system First the capacitive comb drive and the three variations

implemented the transverse drive lateral drive and differential tri-plate drive are

discussed Next the mechanical structure is presented and includes the micro-flexure

selection quality factor and equations of motion Finally instability that places a limit

on the operating range of the sensor as a result of a gap closing capacitive comb drive is

discussed

31 SENSING

Capacitive sensors are based on the parallel plate capacitor and are commonly

implemented using inter-digitated comb fingers A capacitance can be realized if a voltage

is placed between two closely spaced plates If the fringing electric field is neglected the

capacitance of the parallel plate capacitor Cpp (F) is expressed as [28]

air pppp

k AC

dε

= (7)

where kair=1 is the dielectric constant of air at atmospheric pressure ε=88542times10-12 F m-1

is the permittivity of free space App is the overlapping area (m2) and d is the distance

between the two plates (m)

15

16

Figure 3-1 An inter-digitated comb drive where changes in capacitance are generated by either changes in gap distance x0 or in the overlap area y0timest

Figure 3-1

For a set of inter-digitated comb fingers or comb drive shown in the total

capacitance Ctotal is expressed as [28]

0 0

0 0

( ) ( )( ) (Total cy y y yC N t

)x x x xε

+ += + + minus

(8)

where N is the number of comb fingers on the movable comb y0 is the initial comb finger

overlap length (m) tc is the out of plane comb finger thickness (m) and x0 is the inital

finger gap (m) In this work the comb drive is implemented in three arrangements

transverse lateral and differential tri-plate In a transverse comb drive changes in

capacitance are generated by transverse movements along the x-axis resulting in changes

to the gap x0 between the movable and fixed comb fingers The nonlinear relation

between total capacitance CTotalT and displacement is expressed as

17

00 0

1 1( ) (Total TC N ty

)x x x xε

= + minus +

(9)

Although the drive sensitivity part is not constant an average sensitivity can be

assumed for small displacements [38]

Total TC xpart

In a lateral comb drive changes in capacitance are generated by lateral movements

along the y-axis resulting in changes to the overlapping area y0timestc The total capacitance

CTotalL between all comb fingers is expressed as

00

2 ( )Total LN tC yx

yε= +

(10)

The sensitivity of the lateral comb drive SL is expressed as

0

2Total LL

C N tSy x

εpart= =

part

(11)

The lateral comb drive was first introduced by W C Tang T H Nguyen and R T

Howe [33] in an electrostatic resonator and has since been used numerous times in both

drive and sense configurations It has the advantage of a fully linear response and

negligible instability but offers low sensitivity The sensitivity for the lateral drive L1

implemented in this work is 0038 pF microm-1

A differential tri-plate drive shown in Figure 3-2 which is a variation of the

transverse drive is implemented to improve sensitivity and reduce noise [39]

18

Figure 3-2 An inter-digitated comb drive in a differential tri-plate arrangement where x01 x02 and x03 are the gap distances and x01= x02ltlt x03 to improve sensitivity

In the tri-plate drive the at rest gap distances between the two capacitors formed by any

movable finger is not equivalent For a differential arrangement insuring that x01=

x02ltltx03 the linearity for small displacements (less than 1 microm) is enhanced over the case

of a single ended arrangement where x01=x02=x03=x0 which is equivalent to the

transverse drive This is shown graphically in Figure 3-3 where x01= x02=x03=2 microm in

the lower curve and x01= x02= 2 microm and x03= 15 microm in the upper curve A substantial

improvement in linearity is attained with the tri-plate drive

19

0

005

01

015

02

025

03

035

0 02 04 06 08 1

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-3 Comparison of the effect of gap distances on linearity for two case lsquorsquo represents case 1 where the fingers in are implemented in a differential tri-plate arrangement and x01= x02= 2 microm and x03= 15 microm lsquorsquo represents the case 2 where the fingers are implemented in a single ended arrangement and x01= x02=x03=2 microm

The tri-plate drive is also implemented in a differential arrangement where the output

is the change in capacitance rather than the total capacitance resulting from an external

excitation This reduces errors cause by off-axis excitation and common mode noise The

output of the differential tri-plate drive is expressed as [39]

Total Top Total BottomC C C∆ = minus where

003 01

1 1( ) (Total TopC N ty

)x x x xε

= +

minus +

(12)

20

002 03

1 1( ) (Total BottomC N ty

)x x x xε

= +

minus +

A graphic comparison of the capacitancedisplacement relation for sensors T1 and T2 is

shown in Figure 3-4 For these gap closing drives it is assumed that for displacements

less that 1 microm the relation is linear indicated by the dashed line above each curve The

sensitivity within this region is approximately 0046 pF microm-1 and 0253 pF microm-1 for T1

and T2 respectively The order of magnitude improvement is attributed to the differential

tri-plate arrangement and thicker comb finger

0

02

04

06

08

1

12

14

16

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-4 lsquorsquo represents the output of the differential tri-plate drive while lsquorsquo represents the output of the transverse drive The dashed line represents the linear approximation for displacements less that 1 microm For the purpose of comparison the output of the transverse drive was equalized to zero at zero displacement

21

In this work the transverse drive is first implemented in sensor T1 using the

PolyMUMPS fabrication process Then in an effort to linearize the sensor response the

lateral drive is used in L1 using SoiMUMPS The reduction in drive sensitivity due to the

lateral drive is further reduced by increasing the gap distance to 3 microm which is

recommended by the foundry to ensure a successful device These reductions are

balanced by increasing the number of comb fingers N and selecting a substrate thickness

that is 125 times large than T1 The third device T2 is fabricated using SoiMUMPS and

uses the tri-plate differential drive This design capitalizes on the increase in sensitivity of

a gap closing drive and larger substrate thickness attaining an order of magnitude

increase in sensitivity over T1 and L1 The specifications for each drive along with the

successive improvement in drive sensitivity are outlined in Table 3-1

Table 3-1 Specifications for each capacitive comb drive implemented in this work T1 L1 and T2

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Comb Finger Thickness tc (microm) 2 25 25 Gap Distance x0 (microm) 2 3 2 Number of Fingers N 200 256 30 Overlap length y0 (microm) 45 10 70 Drive Sensitivity (pF microm-1) 0046 0038 0273

32 MECHANICAL ANALYSIS

321 MICRO-FLEXURE SELECTION

The sensor structure consists of a proof mass connected to a compliant micro-flexure

that itself is anchored to the sensor substrate The proof mass displaces in response to an

22

applied acceleration and this displacement is sensed using the capacitive comb drive

attached to the proof mass

Proof mass

Folded beam micro-flexure

Anchor

Figure 3-5 Sensor Structure

This section focuses on micro-flexure selection as it should exhibit a large linear

displacement range and have a high lateral stiffness ratio to reduce cross-axis sensitivity

A number of micro-flexures have been used in MEMS research these include the

clamped-clamped crab-leg and folded beam shown in Figure 3-6-through 3-7

The clamped-clamped flexure exhibits a high stiffness ratio however for large lateral

displacements extensional axial forces buildup within the beams resulting in a non-linear

force-displacement relation [40] Within the linear regime the stiffness of the clamped-

clamped flexure is expressed as [40]

3

3b

yEt wk

L=

(13)

23

where E is the Youngrsquos Modulus of the beam material (N m-2) tb is the beam out of

plane thickness (m) w is the beam width (m) and L is the beam length (m) The stiffness

ratio is expressed as [40]

2x

y

k Lk w

=

(14)

If L is two orders of magnitude greater than w this results in a ratio of approximately

10000 The linear displacement range is determined using Finite Element Analysis (FEA)

of a clamped-clamped flexure with L=387 microm w=4 microm and tb=2 microm The result of this

simulation is plotted in and compared to a structure with constant stiffness ky

= 035 N m-1 predicted by Eq (13) The linear relation is valid for small deflections that

are less than 2 microm In addition this micro-flexure is susceptible to buckling as a result of

a compressive residual stress that may arise due to the fabrication process [41]

Figure 3-9

In the crab leg design the thigh section is meant to relieve the extensional axial

stress postponing the occurrence of non-linearity until greater lateral displacement [42]

The stiffness of the crab leg flexure is expressed as [43]

3

44

b s s ty

s s t

Et w L LkL L L

κκ

+= +

(15)

where κ=IsIt=(wswt)3 is a ratio of the second moment of inertia of the shin (s) and thigh

(t) ws and wt represent the width and Ls and Lt represent the length of the shin and thigh

regions respectively The stiffness ratio is expressed as [43]

34

4t s s tx

y s t s

w L L Lkk w L L L

αα t

+= +

(16)

The linear displacement range of a crab leg flexure with Ls=285 microm Lt=30 microm ws=wt=3

microm and tb=2 microm is determined using FEA The results are plotted in Figure 3-9 and

compared to a structure of constant stiffness ky=035 N m-1 predicted by Eq (15) The

24

non-linearity is greatly reduced when compared to the clamp-clamped flexure with Eq

accurately predicting the deflection until 4 microm However the flexure is more

susceptible to off-axis deflection as there is an order of magnitude reduction in the

stiffness ratio as compared to the clamped-clamped flexure

(15)

The folded beam flexure stands out as it is designed to exhibit a large linear

displacement range and to relieve residual stress due to the manufacturing process by

allowing the beams to expand and contract along the axial direction [40] It was first used

in an electrostatic resonator and has subsequently been used for electrostatic actuation in

a magnetic hard drives and micro-positioning [3344] If the truss section is rigid

ItrussgtgtIbeam then the stiffness of the flexure is expressed as [4043]

3

32b

yEt wk

L=

(17)

The stiffness ratio is expressed as [40]

2x

y

k Lk w

=

(18)

Figure 3-6 Clamped-clamped flexure

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

4

14 OPERATING ENVIRONMENT

Through proper packaging MEMS sensors can withstand the harsh automotive

environment where they face cyclic temperatures (-40degC to 150degC) mechanical shock or

vibration (10g-500g) and exposure to various fluids (oil brake fluid or engine coolant)

and airborne particles [8] Temperature changes induce thermal stresses causing

undesirable bending of the silicon microstructures that changes their dynamic response

Mechanical shock which could occur during a high impact collision or vibration from

uneven road conditions cause the delicate microstructures to fracture at points of stress

concentration These effects could be reduced by efficient design Use of a folded beam

micro-flexure that allows expansion and contraction or use of fillets at corners which

reduce the chance of fracture Bonding pads wire bonding and onboard circuit

components can be protected from air borne particles which could cause shorting using

parylene or gels Alternatively the entire package can be sealed hermetically which also

protects the MEMS device [8] In some cases such as fluid pressure sensors intimate

contact between the MEMS device and potentially corrosive fluid is necessary A recent

approach used a nanometer thin coating of silicon carbide (SiC) over a foundry fabricated

MEMS sensor using low temperature low pressure chemical vapor deposition (LPCVD)

[9] The SiC known for its chemical inertness did not compromise the mechanical

characteristics of the original design Therefore MEMS provides a strong platform for an

automotive sensor

5

15 OBJECTIVES AND CONTRIBUTION

The objective of this work is the development of a low frequency MEMS vibration

sensor The sensor must be low cost and mass producible which can be achieved using

standard foundry fabrication processes such as the Polysilicon Multi-User MEMS

Processes (PolyMUMPS) or the Silicon-on-Insulator Multi-User MEMS Processes

(SoiMUMPS) In addition the number of additional processing steps which are required

if integrated CMOS circuitry is used or if vacuum packaging is required must be

minimized This can be achieved using an off-chip circuit made with widely available

discrete components and sensor operation in air

In addition this work presents a MEMS sensor that is fabricated using standard

foundry processes such as PolyMUMPS and SoiMUMPS compared to many MEMS

sensors in literature that are fabricated using customized processes

16 RESEARCH OUTLINE

This work demonstrates three iterations of a low frequency capacitive vibration

sensor fabricated using the standard foundry processes Each sensor consists of a

suspended proof mass that displaces in response to an external vibration The proof mass

is attached to a folded beam micro-flexure compliant in one direction and its lateral

movement is sensed using a capacitive comb drive The first sensor T1 is fabricated in

PolyMUMPS and uses a transverse capacitive comb drive The second L1 fabricated in

SoiMUMPS uses a fully linear lateral capacitive drive and achieves an increase in

sensitivity from T1 Finally the third sensor T2 is fabricated in SoiMUMPS and uses a

differential tri-plate capacitive drive This sensor takes advantage of the increase in

6

sensitivity gained from a transverse drive but rejects off-axis displacement errors and

common mode using a differential tri-plate arrangement In addition an off-chip

capacitance to voltage readout circuit is fabricated and tested on sensor L1

Chapter 2 begins with an overview on the various methods of fault detection

followed by a review of MEMS vibration sensors each with a different transduction

mechanism Chapter 3 outlines the sensor structure focusing on the mechanical design

and sensing principal Chapter 4 focuses on the readout circuitry layout and simulation

results Chapter 5 outlines the various source of noise including mechanical noise arising

from the sensor and electric noise from the readout circuit Chapter 6 overviews the

foundry fabrication processes used including PolyMUMPs and SoiMUMPS The

experimental results of two sensors T1 and L1 which were fabricated and tested within

the thesis time frame are presented in Chapter 7 Chapter 8 concludes the thesis with a

brief summary and outline for future work

2 LITERATURE REVIEW

21 EXISTING METHODS FOR FAULT DETECTION

As this thesis develops a sensor for fault detection using vibration monitoring this

section overviews three existing methods for fault detection in mechanical systems First

motor current monitoring is discussed as it has the ability to measure significant bearing

faults in motor systems Next temperature sensors specifically thin film thermocouples

(TFTC) and MEMS temperature sensors are discussed These sensors are mounted in the

chamber of a lubricating fluid or in close proximity to a moving component and monitor

temperature fluctuations caused by increased wear Finally acoustic emission (AE)

sensors are discussed which have the ability to detect ultrasonic emissions generated by

plastic deformation in the crystalline lattice of a mechanical component

211 MOTOR CURRENT MONITORING

Monitoring of stator current is used as a method of determining bearing related faults

that produce radial motion between rotor and stator This motion varies the air gap flux

density producing stator currents at predictable frequencies fbng that represent the

bearing faults [10]

bng e vf f m f= plusmn sdot (2)

where fe is the electrical supply frequency m=123hellip and fv is one of the characteristics

frequencies listed in the equations of (1)The bearing damage must be significant to

produce noticeable current fluctuations and it may be difficult to distinguish from current

fluctuations due to voltage imbalances machine saturation or statorarmature faults [2]

7

8

212 TEMPERATURE SENSORS

Temperature sensors are largely used in high temperature high load and high speed

environments typically turbine engines and industrial tools Thermocouple sensors

consist of two dissimilar metals joined together at two junctions and when exposed to a

temperature difference generate a thermoelectric voltage dependant only on the material

properties and junction temperature difference [11] In turbine engines these sensors can

be placed on bearing casings or in the lubricant to monitor temperature fluctuations due

to increased friction The thermocouples are fabricated using bulk materials or micro-

machined thin films (TFTC) with the latter providing a faster response small package

and substantially lower production costs [12] The TFTC can be embedded for in situ

monitoring of processes in hostile environments One study successfully embedded

TFTC into electroplated nickel and attained device characteristics on par with surface

mounted devices [13] More recently MEMS temperature sensors have been developed

that consist of micro-machined semiconductor material that undergoes structural

deformation with temperature changes A common implementation is a bent beam

structure whose temperature induced deflection δ is expressed as [14]

( ) 3

2 2

sin12 cos sinz

EA T LEI AEL

α βδ 2β β

∆=

+

(3)

Temperature sensors are well suited for localized fault monitoring and detection but can

only characterize a mechanical system consisting of multiple components through the use

of multiple sensors

9

213 ACOUSTIC EMISSION SENSORS

Acoustic emission (AE) sensors have been used to characterize wear in machine

tools and monitor bearing and gear problems in centrifugal pumps [1516] First

developed as a Non-Destructive Testing (NDT) technique to detect cracks in civil

structures these sensors detect acoustic emissions generated by the release of vibration

waves in a crystalline lattice due to plastic deformation or crack growth [16]

Measurements are made using piezoelectric transducers with high natural frequencies

100 kHz to 1 MHz to capture the ultrasonic AE emissions An AE sensor is useful as it

has the ability to detect subsurface cracks in gear teeth or bearings before appearing on

the surface causing further damage [17] More recently MEMS acoustic sensors have

been developed and one design includes multiple transducers on a single substrate

which each detect acoustic emission energy at different frequencies [18] This helps

distinguish spurious acoustic emissions arising from impact and friction from those

arising from plastic deformation When compared to typical piezoelectric sensors the

MEMS devices have lower sensitivities and fail to detect some acoustic emissions

[1819] In addition the acoustic emission signal suffers severe attenuation as is crosses

various interfaces such as a gearbox or bearing casings In one experiment consisting of

a pinion gear and an associated bearing a 44dB attenuation was seen between an AE

sensor placed directly on the pinion to one placed on the bearing casing and in some

cases intermediate loss of the signal was observed [20]

10

22 MEMS VIBRATION SENSORS

This thesis focuses on the development of a MEMS vibration sensor for fault

detection as the vibration signature of a mechanical system has the potential to give an

overview of the entire system Many types of MEMS vibration sensors exist with

different operating principals based on resonance the piezoelectric effect and

displacement variation

221 RESONANT SENSORS

Resonant sensors are based on the resonant frequency shift of a fixed-fixed micro-

beam that is excited into resonance using electrostatic thermal or piezoelectric methods

Under an applied axial load there is a shift in the resonant frequency which is expressed

as [21]

2 2

2

1 12 1

nn n

EI Fl2

fl A Eα γ

π ρ= +

I

(4)

where l is the beam length E is Youngrsquos Modulus I is the moment of inertia ρ is the

density A is the cross section area F is the applied axial force αn and γn are mode

dependant coefficients and are 4730 and 0295 for mode 1 in the case discussed The

axial force F is generated by a proof mass attached to one beam end which moves in

response to an external acceleration The performance of resonant sensors are dependant

on the quality factor Q which is the ratio of total energy stored in the system to the

energy lost per cycle and determines the sharpness and amplitude of the resonance peak

A high Q resonator develops a peak that is easily distinguishable in a phase locked loop

control circuit and allows for improved performance and resolution [3] A resonator Q

value is lowered by energy losses to the surrounding fluid attached supports and

11

material Losses to the fluid surrounding the MEMS structure is the dominant loss

mechanism for MEMS sensors operated in air and can be greatly reduced by vacuum

packaging which can be accomplished using glass-silicon anodic bonding [22] In

addition to increasing fabrication cost the long term stability of the vacuum degrades

with ultimate device failure stemming from leakage through micro cracks and defects

and out-gassing [22] Support losses arise from the restoring forces they generate and

can be reduce by balanced designs such as a double ended tuning fork (DETF) or

operation in higher modes [3] In addition the sensitivity of a resonant sensor can be

increased by the addition of micro-levers to increase the axial force In one such design a

two stage micro-lever increased the sensitivity by an order of magnitude over a more

conventional single lever arm [2324]

222 PIEZOELECTRIC SENSORS

Piezoelectric materials produce a charge in response to an applied force This is an

intrinsic effect in materials such as Zinc Oxide (ZnO) whereas materials such as Barium

Titanate or Lead Zirconate Titanate (PZT) need to be poled by placing the material in a

strong electric field at an elevated temperature [3] The piezoelectric effect is expressed

mathematical by [25]

i ij j ik kD d T Eε= + (5)

where D represents the electrical displacement (Cm2) d is the charge coefficient (CN)

T is the applied stress matrix (Nm2) ε is the permittivity matrix (Fm) and E is the

electric field matrix (Vm) The subscripts i=123 j=123456 and k=123 represent

the direction to which the physical property is related For micro-sensors poled in the out

of plane direction only the D3 term is relevant and if fabricated using ZnO or PZT only

12

the charge coefficients d31 d32 and d33 are relevant [25] The piezoelectric material is

commonly deposited as a thin film on the surface of compliant structures in bulk micro-

machined sensors or incorporated as a layer in surface micro-machining on a similar

structure [26] However the inclusion of multiple materials introduces large temperature

variations causing out of plane bending especially in surface micro-machined devices

[2627] Poled piezoelectric materials could be depolarized which serves to reduce the

piezoelectric affect in the material This can occur if the sensor is exposed to a strong

electric field of the opposite polarity high temperatures in excess of the Curie point or

high mechanical stress Also the piezoelectric effect reduces over time an effect that can

be reduced with the addition of composite elements [28] Foundry processes such as

PolyMUMPS and SoiMUMPS do not incorporate piezoelectric layers thus sensors using

this method of transduction must use a customized process The piezoelectric sensor

outlined in [26] uses polysilicon surface micro-machining similar to PolyMUMPS

however it incorporates a ZnO layer that is not available in the foundry process

223 DISPLACEMENT VARIATION SENSORS

Displacement variation sensors consist of a proof mass connected to a compliant

micro-flexure The movement of the proof mass is sensed using optical electron

tunneling or more commonly capacitive methods A recent optical accelerometer

utilized a novel multilayer nano-grating for in-plane displacement sensing of the proof

mass [29] For small space variations between the nano-gratings a large change in the

optical reflection amplitude of the grating was observed [30] Even though optical

sensors offer resolutions approaching the Brownian noise limit they are not available in

small packages that are easily placed in space constricted areas for a low cost [29]

13

The electron tunneling effect is observed when the proximity between two electrodes

one flexible and the other fixed is on the order of 10Ǻ [31] If the two electrodes are

biased with voltage VB a tunneling current It flows between the gap [31]

exp( )t B I tgI V xαinfin sdot minus Φ (6)

where α1 is a constant and equal to 1025 Ǻ-1 eV-frac12 Ф is the effective height of the

tunneling barrier and xtg is the tunneling gap When the flexible electrode moves in

response to acceleration a feedback controller maintains the original distance while

determining the applied acceleration Electron tunneling sensors are able to sense

accelerations in the microg-ng range due to the exponential relation between tunneling current

and gap However these highly sensitive devices are susceptible to thermo-mechanical

noise which could be reduced by operation at low pressure [32] It is also important for

the two electrodes to be coated with metal which is difficult to achieve with foundry

fabrication

Capacitive sensors are based on the parallel plate capacitor and are implemented

using inter-digitated comb fingers in an in-plane lateral or transverse drive or in an out of

plane parallel plate drive [33] These sensors have the advantage of high sensitivity low

power consumption small package low temperature dependence and are easily

integrated with existing foundry fabrication processes [34] The lateral drive offers a fully

linear response but low sensitivity whereas the transverse and parallel plate drives

achieve high sensitivity if gap distances are of a few microns and large capacitance areas

are utilized However both drives have a non-linear response and are prone to pull-in

Also if these sensors are fabricated using a surface micromachining process such as

PolyMUMPS where the ground plane is a few microns above the structural layer they

14

are prone to electrical failure due to shorting This was observed in an experiment where

similar sensors were placed in a vibrating environment with a peak of 120g and shorting

between the structural layer and ground plane was found to be the predominate mode of

failure [35] However with certain fabrication processes such as SoiMUMPS this issue

can be neglected as the substrate material below any moving component is removed

Most capacitance sensors are fabricated using an integrated CMOS (Complementary

Metal Oxide Semiconductor)-MEMS technique where the MEMS sensor is fabricated

before during or after the CMOS circuit fabrication [36] Although these sensors offer

high-sensitivity and small parasitic capacitance they suffer from in-plane and out of

plane curling of the beam which reduces the capacitance between adjacent comb fingers

[37] In addition multiple processing steps are required for device fabrication

More than the other transduction methods MEMS capacitive displacement sensors are

attractive for practical implementation in an automobile and therefore is the selected

method for the vibration sensor presented in this work Their low temperature dependence

is an ideal characteristic for an automobile environment that faces varying temperature If

implemented with off-chip readout circuitry they require a standard foundry fabrication

process which provides a strong platform for mass production and low unit cost

3 SENSOR STRUCTURE

This chapter presents the physical structure of the three MEMS vibration sensors

which includes the displacement sensing capacitive comb drive and the micro-

flexureproof mass system First the capacitive comb drive and the three variations

implemented the transverse drive lateral drive and differential tri-plate drive are

discussed Next the mechanical structure is presented and includes the micro-flexure

selection quality factor and equations of motion Finally instability that places a limit

on the operating range of the sensor as a result of a gap closing capacitive comb drive is

discussed

31 SENSING

Capacitive sensors are based on the parallel plate capacitor and are commonly

implemented using inter-digitated comb fingers A capacitance can be realized if a voltage

is placed between two closely spaced plates If the fringing electric field is neglected the

capacitance of the parallel plate capacitor Cpp (F) is expressed as [28]

air pppp

k AC

dε

= (7)

where kair=1 is the dielectric constant of air at atmospheric pressure ε=88542times10-12 F m-1

is the permittivity of free space App is the overlapping area (m2) and d is the distance

between the two plates (m)

15

16

Figure 3-1 An inter-digitated comb drive where changes in capacitance are generated by either changes in gap distance x0 or in the overlap area y0timest

Figure 3-1

For a set of inter-digitated comb fingers or comb drive shown in the total

capacitance Ctotal is expressed as [28]

0 0

0 0

( ) ( )( ) (Total cy y y yC N t

)x x x xε

+ += + + minus

(8)

where N is the number of comb fingers on the movable comb y0 is the initial comb finger

overlap length (m) tc is the out of plane comb finger thickness (m) and x0 is the inital

finger gap (m) In this work the comb drive is implemented in three arrangements

transverse lateral and differential tri-plate In a transverse comb drive changes in

capacitance are generated by transverse movements along the x-axis resulting in changes

to the gap x0 between the movable and fixed comb fingers The nonlinear relation

between total capacitance CTotalT and displacement is expressed as

17

00 0

1 1( ) (Total TC N ty

)x x x xε

= + minus +

(9)

Although the drive sensitivity part is not constant an average sensitivity can be

assumed for small displacements [38]

Total TC xpart

In a lateral comb drive changes in capacitance are generated by lateral movements

along the y-axis resulting in changes to the overlapping area y0timestc The total capacitance

CTotalL between all comb fingers is expressed as

00

2 ( )Total LN tC yx

yε= +

(10)

The sensitivity of the lateral comb drive SL is expressed as

0

2Total LL

C N tSy x

εpart= =

part

(11)

The lateral comb drive was first introduced by W C Tang T H Nguyen and R T

Howe [33] in an electrostatic resonator and has since been used numerous times in both

drive and sense configurations It has the advantage of a fully linear response and

negligible instability but offers low sensitivity The sensitivity for the lateral drive L1

implemented in this work is 0038 pF microm-1

A differential tri-plate drive shown in Figure 3-2 which is a variation of the

transverse drive is implemented to improve sensitivity and reduce noise [39]

18

Figure 3-2 An inter-digitated comb drive in a differential tri-plate arrangement where x01 x02 and x03 are the gap distances and x01= x02ltlt x03 to improve sensitivity

In the tri-plate drive the at rest gap distances between the two capacitors formed by any

movable finger is not equivalent For a differential arrangement insuring that x01=

x02ltltx03 the linearity for small displacements (less than 1 microm) is enhanced over the case

of a single ended arrangement where x01=x02=x03=x0 which is equivalent to the

transverse drive This is shown graphically in Figure 3-3 where x01= x02=x03=2 microm in

the lower curve and x01= x02= 2 microm and x03= 15 microm in the upper curve A substantial

improvement in linearity is attained with the tri-plate drive

19

0

005

01

015

02

025

03

035

0 02 04 06 08 1

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-3 Comparison of the effect of gap distances on linearity for two case lsquorsquo represents case 1 where the fingers in are implemented in a differential tri-plate arrangement and x01= x02= 2 microm and x03= 15 microm lsquorsquo represents the case 2 where the fingers are implemented in a single ended arrangement and x01= x02=x03=2 microm

The tri-plate drive is also implemented in a differential arrangement where the output

is the change in capacitance rather than the total capacitance resulting from an external

excitation This reduces errors cause by off-axis excitation and common mode noise The

output of the differential tri-plate drive is expressed as [39]

Total Top Total BottomC C C∆ = minus where

003 01

1 1( ) (Total TopC N ty

)x x x xε

= +

minus +

(12)

20

002 03

1 1( ) (Total BottomC N ty

)x x x xε

= +

minus +

A graphic comparison of the capacitancedisplacement relation for sensors T1 and T2 is

shown in Figure 3-4 For these gap closing drives it is assumed that for displacements

less that 1 microm the relation is linear indicated by the dashed line above each curve The

sensitivity within this region is approximately 0046 pF microm-1 and 0253 pF microm-1 for T1

and T2 respectively The order of magnitude improvement is attributed to the differential

tri-plate arrangement and thicker comb finger

0

02

04

06

08

1

12

14

16

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-4 lsquorsquo represents the output of the differential tri-plate drive while lsquorsquo represents the output of the transverse drive The dashed line represents the linear approximation for displacements less that 1 microm For the purpose of comparison the output of the transverse drive was equalized to zero at zero displacement

21

In this work the transverse drive is first implemented in sensor T1 using the

PolyMUMPS fabrication process Then in an effort to linearize the sensor response the

lateral drive is used in L1 using SoiMUMPS The reduction in drive sensitivity due to the

lateral drive is further reduced by increasing the gap distance to 3 microm which is

recommended by the foundry to ensure a successful device These reductions are

balanced by increasing the number of comb fingers N and selecting a substrate thickness

that is 125 times large than T1 The third device T2 is fabricated using SoiMUMPS and

uses the tri-plate differential drive This design capitalizes on the increase in sensitivity of

a gap closing drive and larger substrate thickness attaining an order of magnitude

increase in sensitivity over T1 and L1 The specifications for each drive along with the

successive improvement in drive sensitivity are outlined in Table 3-1

Table 3-1 Specifications for each capacitive comb drive implemented in this work T1 L1 and T2

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Comb Finger Thickness tc (microm) 2 25 25 Gap Distance x0 (microm) 2 3 2 Number of Fingers N 200 256 30 Overlap length y0 (microm) 45 10 70 Drive Sensitivity (pF microm-1) 0046 0038 0273

32 MECHANICAL ANALYSIS

321 MICRO-FLEXURE SELECTION

The sensor structure consists of a proof mass connected to a compliant micro-flexure

that itself is anchored to the sensor substrate The proof mass displaces in response to an

22

applied acceleration and this displacement is sensed using the capacitive comb drive

attached to the proof mass

Proof mass

Folded beam micro-flexure

Anchor

Figure 3-5 Sensor Structure

This section focuses on micro-flexure selection as it should exhibit a large linear

displacement range and have a high lateral stiffness ratio to reduce cross-axis sensitivity

A number of micro-flexures have been used in MEMS research these include the

clamped-clamped crab-leg and folded beam shown in Figure 3-6-through 3-7

The clamped-clamped flexure exhibits a high stiffness ratio however for large lateral

displacements extensional axial forces buildup within the beams resulting in a non-linear

force-displacement relation [40] Within the linear regime the stiffness of the clamped-

clamped flexure is expressed as [40]

3

3b

yEt wk

L=

(13)

23

where E is the Youngrsquos Modulus of the beam material (N m-2) tb is the beam out of

plane thickness (m) w is the beam width (m) and L is the beam length (m) The stiffness

ratio is expressed as [40]

2x

y

k Lk w

=

(14)

If L is two orders of magnitude greater than w this results in a ratio of approximately

10000 The linear displacement range is determined using Finite Element Analysis (FEA)

of a clamped-clamped flexure with L=387 microm w=4 microm and tb=2 microm The result of this

simulation is plotted in and compared to a structure with constant stiffness ky

= 035 N m-1 predicted by Eq (13) The linear relation is valid for small deflections that

are less than 2 microm In addition this micro-flexure is susceptible to buckling as a result of

a compressive residual stress that may arise due to the fabrication process [41]

Figure 3-9

In the crab leg design the thigh section is meant to relieve the extensional axial

stress postponing the occurrence of non-linearity until greater lateral displacement [42]

The stiffness of the crab leg flexure is expressed as [43]

3

44

b s s ty

s s t

Et w L LkL L L

κκ

+= +

(15)

where κ=IsIt=(wswt)3 is a ratio of the second moment of inertia of the shin (s) and thigh

(t) ws and wt represent the width and Ls and Lt represent the length of the shin and thigh

regions respectively The stiffness ratio is expressed as [43]

34

4t s s tx

y s t s

w L L Lkk w L L L

αα t

+= +

(16)

The linear displacement range of a crab leg flexure with Ls=285 microm Lt=30 microm ws=wt=3

microm and tb=2 microm is determined using FEA The results are plotted in Figure 3-9 and

compared to a structure of constant stiffness ky=035 N m-1 predicted by Eq (15) The

24

non-linearity is greatly reduced when compared to the clamp-clamped flexure with Eq

accurately predicting the deflection until 4 microm However the flexure is more

susceptible to off-axis deflection as there is an order of magnitude reduction in the

stiffness ratio as compared to the clamped-clamped flexure

(15)

The folded beam flexure stands out as it is designed to exhibit a large linear

displacement range and to relieve residual stress due to the manufacturing process by

allowing the beams to expand and contract along the axial direction [40] It was first used

in an electrostatic resonator and has subsequently been used for electrostatic actuation in

a magnetic hard drives and micro-positioning [3344] If the truss section is rigid

ItrussgtgtIbeam then the stiffness of the flexure is expressed as [4043]

3

32b

yEt wk

L=

(17)

The stiffness ratio is expressed as [40]

2x

y

k Lk w

=

(18)

Figure 3-6 Clamped-clamped flexure

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

5

15 OBJECTIVES AND CONTRIBUTION

The objective of this work is the development of a low frequency MEMS vibration

sensor The sensor must be low cost and mass producible which can be achieved using

standard foundry fabrication processes such as the Polysilicon Multi-User MEMS

Processes (PolyMUMPS) or the Silicon-on-Insulator Multi-User MEMS Processes

(SoiMUMPS) In addition the number of additional processing steps which are required

if integrated CMOS circuitry is used or if vacuum packaging is required must be

minimized This can be achieved using an off-chip circuit made with widely available

discrete components and sensor operation in air

In addition this work presents a MEMS sensor that is fabricated using standard

foundry processes such as PolyMUMPS and SoiMUMPS compared to many MEMS

sensors in literature that are fabricated using customized processes

16 RESEARCH OUTLINE

This work demonstrates three iterations of a low frequency capacitive vibration

sensor fabricated using the standard foundry processes Each sensor consists of a

suspended proof mass that displaces in response to an external vibration The proof mass

is attached to a folded beam micro-flexure compliant in one direction and its lateral

movement is sensed using a capacitive comb drive The first sensor T1 is fabricated in

PolyMUMPS and uses a transverse capacitive comb drive The second L1 fabricated in

SoiMUMPS uses a fully linear lateral capacitive drive and achieves an increase in

sensitivity from T1 Finally the third sensor T2 is fabricated in SoiMUMPS and uses a

differential tri-plate capacitive drive This sensor takes advantage of the increase in

6

sensitivity gained from a transverse drive but rejects off-axis displacement errors and

common mode using a differential tri-plate arrangement In addition an off-chip

capacitance to voltage readout circuit is fabricated and tested on sensor L1

Chapter 2 begins with an overview on the various methods of fault detection

followed by a review of MEMS vibration sensors each with a different transduction

mechanism Chapter 3 outlines the sensor structure focusing on the mechanical design

and sensing principal Chapter 4 focuses on the readout circuitry layout and simulation

results Chapter 5 outlines the various source of noise including mechanical noise arising

from the sensor and electric noise from the readout circuit Chapter 6 overviews the

foundry fabrication processes used including PolyMUMPs and SoiMUMPS The

experimental results of two sensors T1 and L1 which were fabricated and tested within

the thesis time frame are presented in Chapter 7 Chapter 8 concludes the thesis with a

brief summary and outline for future work

2 LITERATURE REVIEW

21 EXISTING METHODS FOR FAULT DETECTION

As this thesis develops a sensor for fault detection using vibration monitoring this

section overviews three existing methods for fault detection in mechanical systems First

motor current monitoring is discussed as it has the ability to measure significant bearing

faults in motor systems Next temperature sensors specifically thin film thermocouples

(TFTC) and MEMS temperature sensors are discussed These sensors are mounted in the

chamber of a lubricating fluid or in close proximity to a moving component and monitor

temperature fluctuations caused by increased wear Finally acoustic emission (AE)

sensors are discussed which have the ability to detect ultrasonic emissions generated by

plastic deformation in the crystalline lattice of a mechanical component

211 MOTOR CURRENT MONITORING

Monitoring of stator current is used as a method of determining bearing related faults

that produce radial motion between rotor and stator This motion varies the air gap flux

density producing stator currents at predictable frequencies fbng that represent the

bearing faults [10]

bng e vf f m f= plusmn sdot (2)

where fe is the electrical supply frequency m=123hellip and fv is one of the characteristics

frequencies listed in the equations of (1)The bearing damage must be significant to

produce noticeable current fluctuations and it may be difficult to distinguish from current

fluctuations due to voltage imbalances machine saturation or statorarmature faults [2]

7

8

212 TEMPERATURE SENSORS

Temperature sensors are largely used in high temperature high load and high speed

environments typically turbine engines and industrial tools Thermocouple sensors

consist of two dissimilar metals joined together at two junctions and when exposed to a

temperature difference generate a thermoelectric voltage dependant only on the material

properties and junction temperature difference [11] In turbine engines these sensors can

be placed on bearing casings or in the lubricant to monitor temperature fluctuations due

to increased friction The thermocouples are fabricated using bulk materials or micro-

machined thin films (TFTC) with the latter providing a faster response small package

and substantially lower production costs [12] The TFTC can be embedded for in situ

monitoring of processes in hostile environments One study successfully embedded

TFTC into electroplated nickel and attained device characteristics on par with surface

mounted devices [13] More recently MEMS temperature sensors have been developed

that consist of micro-machined semiconductor material that undergoes structural

deformation with temperature changes A common implementation is a bent beam

structure whose temperature induced deflection δ is expressed as [14]

( ) 3

2 2

sin12 cos sinz

EA T LEI AEL

α βδ 2β β

∆=

+

(3)

Temperature sensors are well suited for localized fault monitoring and detection but can

only characterize a mechanical system consisting of multiple components through the use

of multiple sensors

9

213 ACOUSTIC EMISSION SENSORS

Acoustic emission (AE) sensors have been used to characterize wear in machine

tools and monitor bearing and gear problems in centrifugal pumps [1516] First

developed as a Non-Destructive Testing (NDT) technique to detect cracks in civil

structures these sensors detect acoustic emissions generated by the release of vibration

waves in a crystalline lattice due to plastic deformation or crack growth [16]

Measurements are made using piezoelectric transducers with high natural frequencies

100 kHz to 1 MHz to capture the ultrasonic AE emissions An AE sensor is useful as it

has the ability to detect subsurface cracks in gear teeth or bearings before appearing on

the surface causing further damage [17] More recently MEMS acoustic sensors have

been developed and one design includes multiple transducers on a single substrate

which each detect acoustic emission energy at different frequencies [18] This helps

distinguish spurious acoustic emissions arising from impact and friction from those

arising from plastic deformation When compared to typical piezoelectric sensors the

MEMS devices have lower sensitivities and fail to detect some acoustic emissions

[1819] In addition the acoustic emission signal suffers severe attenuation as is crosses

various interfaces such as a gearbox or bearing casings In one experiment consisting of

a pinion gear and an associated bearing a 44dB attenuation was seen between an AE

sensor placed directly on the pinion to one placed on the bearing casing and in some

cases intermediate loss of the signal was observed [20]

10

22 MEMS VIBRATION SENSORS

This thesis focuses on the development of a MEMS vibration sensor for fault

detection as the vibration signature of a mechanical system has the potential to give an

overview of the entire system Many types of MEMS vibration sensors exist with

different operating principals based on resonance the piezoelectric effect and

displacement variation

221 RESONANT SENSORS

Resonant sensors are based on the resonant frequency shift of a fixed-fixed micro-

beam that is excited into resonance using electrostatic thermal or piezoelectric methods

Under an applied axial load there is a shift in the resonant frequency which is expressed

as [21]

2 2

2

1 12 1

nn n

EI Fl2

fl A Eα γ

π ρ= +

I

(4)

where l is the beam length E is Youngrsquos Modulus I is the moment of inertia ρ is the

density A is the cross section area F is the applied axial force αn and γn are mode

dependant coefficients and are 4730 and 0295 for mode 1 in the case discussed The

axial force F is generated by a proof mass attached to one beam end which moves in

response to an external acceleration The performance of resonant sensors are dependant

on the quality factor Q which is the ratio of total energy stored in the system to the

energy lost per cycle and determines the sharpness and amplitude of the resonance peak

A high Q resonator develops a peak that is easily distinguishable in a phase locked loop

control circuit and allows for improved performance and resolution [3] A resonator Q

value is lowered by energy losses to the surrounding fluid attached supports and

11

material Losses to the fluid surrounding the MEMS structure is the dominant loss

mechanism for MEMS sensors operated in air and can be greatly reduced by vacuum

packaging which can be accomplished using glass-silicon anodic bonding [22] In

addition to increasing fabrication cost the long term stability of the vacuum degrades

with ultimate device failure stemming from leakage through micro cracks and defects

and out-gassing [22] Support losses arise from the restoring forces they generate and

can be reduce by balanced designs such as a double ended tuning fork (DETF) or

operation in higher modes [3] In addition the sensitivity of a resonant sensor can be

increased by the addition of micro-levers to increase the axial force In one such design a

two stage micro-lever increased the sensitivity by an order of magnitude over a more

conventional single lever arm [2324]

222 PIEZOELECTRIC SENSORS

Piezoelectric materials produce a charge in response to an applied force This is an

intrinsic effect in materials such as Zinc Oxide (ZnO) whereas materials such as Barium

Titanate or Lead Zirconate Titanate (PZT) need to be poled by placing the material in a

strong electric field at an elevated temperature [3] The piezoelectric effect is expressed

mathematical by [25]

i ij j ik kD d T Eε= + (5)

where D represents the electrical displacement (Cm2) d is the charge coefficient (CN)

T is the applied stress matrix (Nm2) ε is the permittivity matrix (Fm) and E is the

electric field matrix (Vm) The subscripts i=123 j=123456 and k=123 represent

the direction to which the physical property is related For micro-sensors poled in the out

of plane direction only the D3 term is relevant and if fabricated using ZnO or PZT only

12

the charge coefficients d31 d32 and d33 are relevant [25] The piezoelectric material is

commonly deposited as a thin film on the surface of compliant structures in bulk micro-

machined sensors or incorporated as a layer in surface micro-machining on a similar

structure [26] However the inclusion of multiple materials introduces large temperature

variations causing out of plane bending especially in surface micro-machined devices

[2627] Poled piezoelectric materials could be depolarized which serves to reduce the

piezoelectric affect in the material This can occur if the sensor is exposed to a strong

electric field of the opposite polarity high temperatures in excess of the Curie point or

high mechanical stress Also the piezoelectric effect reduces over time an effect that can

be reduced with the addition of composite elements [28] Foundry processes such as

PolyMUMPS and SoiMUMPS do not incorporate piezoelectric layers thus sensors using

this method of transduction must use a customized process The piezoelectric sensor

outlined in [26] uses polysilicon surface micro-machining similar to PolyMUMPS

however it incorporates a ZnO layer that is not available in the foundry process

223 DISPLACEMENT VARIATION SENSORS

Displacement variation sensors consist of a proof mass connected to a compliant

micro-flexure The movement of the proof mass is sensed using optical electron

tunneling or more commonly capacitive methods A recent optical accelerometer

utilized a novel multilayer nano-grating for in-plane displacement sensing of the proof

mass [29] For small space variations between the nano-gratings a large change in the

optical reflection amplitude of the grating was observed [30] Even though optical

sensors offer resolutions approaching the Brownian noise limit they are not available in

small packages that are easily placed in space constricted areas for a low cost [29]

13

The electron tunneling effect is observed when the proximity between two electrodes

one flexible and the other fixed is on the order of 10Ǻ [31] If the two electrodes are

biased with voltage VB a tunneling current It flows between the gap [31]

exp( )t B I tgI V xαinfin sdot minus Φ (6)

where α1 is a constant and equal to 1025 Ǻ-1 eV-frac12 Ф is the effective height of the

tunneling barrier and xtg is the tunneling gap When the flexible electrode moves in

response to acceleration a feedback controller maintains the original distance while

determining the applied acceleration Electron tunneling sensors are able to sense

accelerations in the microg-ng range due to the exponential relation between tunneling current

and gap However these highly sensitive devices are susceptible to thermo-mechanical

noise which could be reduced by operation at low pressure [32] It is also important for

the two electrodes to be coated with metal which is difficult to achieve with foundry

fabrication

Capacitive sensors are based on the parallel plate capacitor and are implemented

using inter-digitated comb fingers in an in-plane lateral or transverse drive or in an out of

plane parallel plate drive [33] These sensors have the advantage of high sensitivity low

power consumption small package low temperature dependence and are easily

integrated with existing foundry fabrication processes [34] The lateral drive offers a fully

linear response but low sensitivity whereas the transverse and parallel plate drives

achieve high sensitivity if gap distances are of a few microns and large capacitance areas

are utilized However both drives have a non-linear response and are prone to pull-in

Also if these sensors are fabricated using a surface micromachining process such as

PolyMUMPS where the ground plane is a few microns above the structural layer they

14

are prone to electrical failure due to shorting This was observed in an experiment where

similar sensors were placed in a vibrating environment with a peak of 120g and shorting

between the structural layer and ground plane was found to be the predominate mode of

failure [35] However with certain fabrication processes such as SoiMUMPS this issue

can be neglected as the substrate material below any moving component is removed

Most capacitance sensors are fabricated using an integrated CMOS (Complementary

Metal Oxide Semiconductor)-MEMS technique where the MEMS sensor is fabricated

before during or after the CMOS circuit fabrication [36] Although these sensors offer

high-sensitivity and small parasitic capacitance they suffer from in-plane and out of

plane curling of the beam which reduces the capacitance between adjacent comb fingers

[37] In addition multiple processing steps are required for device fabrication

More than the other transduction methods MEMS capacitive displacement sensors are

attractive for practical implementation in an automobile and therefore is the selected

method for the vibration sensor presented in this work Their low temperature dependence

is an ideal characteristic for an automobile environment that faces varying temperature If

implemented with off-chip readout circuitry they require a standard foundry fabrication

process which provides a strong platform for mass production and low unit cost

3 SENSOR STRUCTURE

This chapter presents the physical structure of the three MEMS vibration sensors

which includes the displacement sensing capacitive comb drive and the micro-

flexureproof mass system First the capacitive comb drive and the three variations

implemented the transverse drive lateral drive and differential tri-plate drive are

discussed Next the mechanical structure is presented and includes the micro-flexure

selection quality factor and equations of motion Finally instability that places a limit

on the operating range of the sensor as a result of a gap closing capacitive comb drive is

discussed

31 SENSING

Capacitive sensors are based on the parallel plate capacitor and are commonly

implemented using inter-digitated comb fingers A capacitance can be realized if a voltage

is placed between two closely spaced plates If the fringing electric field is neglected the

capacitance of the parallel plate capacitor Cpp (F) is expressed as [28]

air pppp

k AC

dε

= (7)

where kair=1 is the dielectric constant of air at atmospheric pressure ε=88542times10-12 F m-1

is the permittivity of free space App is the overlapping area (m2) and d is the distance

between the two plates (m)

15

16

Figure 3-1 An inter-digitated comb drive where changes in capacitance are generated by either changes in gap distance x0 or in the overlap area y0timest

Figure 3-1

For a set of inter-digitated comb fingers or comb drive shown in the total

capacitance Ctotal is expressed as [28]

0 0

0 0

( ) ( )( ) (Total cy y y yC N t

)x x x xε

+ += + + minus

(8)

where N is the number of comb fingers on the movable comb y0 is the initial comb finger

overlap length (m) tc is the out of plane comb finger thickness (m) and x0 is the inital

finger gap (m) In this work the comb drive is implemented in three arrangements

transverse lateral and differential tri-plate In a transverse comb drive changes in

capacitance are generated by transverse movements along the x-axis resulting in changes

to the gap x0 between the movable and fixed comb fingers The nonlinear relation

between total capacitance CTotalT and displacement is expressed as

17

00 0

1 1( ) (Total TC N ty

)x x x xε

= + minus +

(9)

Although the drive sensitivity part is not constant an average sensitivity can be

assumed for small displacements [38]

Total TC xpart

In a lateral comb drive changes in capacitance are generated by lateral movements

along the y-axis resulting in changes to the overlapping area y0timestc The total capacitance

CTotalL between all comb fingers is expressed as

00

2 ( )Total LN tC yx

yε= +

(10)

The sensitivity of the lateral comb drive SL is expressed as

0

2Total LL

C N tSy x

εpart= =

part

(11)

The lateral comb drive was first introduced by W C Tang T H Nguyen and R T

Howe [33] in an electrostatic resonator and has since been used numerous times in both

drive and sense configurations It has the advantage of a fully linear response and

negligible instability but offers low sensitivity The sensitivity for the lateral drive L1

implemented in this work is 0038 pF microm-1

A differential tri-plate drive shown in Figure 3-2 which is a variation of the

transverse drive is implemented to improve sensitivity and reduce noise [39]

18

Figure 3-2 An inter-digitated comb drive in a differential tri-plate arrangement where x01 x02 and x03 are the gap distances and x01= x02ltlt x03 to improve sensitivity

In the tri-plate drive the at rest gap distances between the two capacitors formed by any

movable finger is not equivalent For a differential arrangement insuring that x01=

x02ltltx03 the linearity for small displacements (less than 1 microm) is enhanced over the case

of a single ended arrangement where x01=x02=x03=x0 which is equivalent to the

transverse drive This is shown graphically in Figure 3-3 where x01= x02=x03=2 microm in

the lower curve and x01= x02= 2 microm and x03= 15 microm in the upper curve A substantial

improvement in linearity is attained with the tri-plate drive

19

0

005

01

015

02

025

03

035

0 02 04 06 08 1

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-3 Comparison of the effect of gap distances on linearity for two case lsquorsquo represents case 1 where the fingers in are implemented in a differential tri-plate arrangement and x01= x02= 2 microm and x03= 15 microm lsquorsquo represents the case 2 where the fingers are implemented in a single ended arrangement and x01= x02=x03=2 microm

The tri-plate drive is also implemented in a differential arrangement where the output

is the change in capacitance rather than the total capacitance resulting from an external

excitation This reduces errors cause by off-axis excitation and common mode noise The

output of the differential tri-plate drive is expressed as [39]

Total Top Total BottomC C C∆ = minus where

003 01

1 1( ) (Total TopC N ty

)x x x xε

= +

minus +

(12)

20

002 03

1 1( ) (Total BottomC N ty

)x x x xε

= +

minus +

A graphic comparison of the capacitancedisplacement relation for sensors T1 and T2 is

shown in Figure 3-4 For these gap closing drives it is assumed that for displacements

less that 1 microm the relation is linear indicated by the dashed line above each curve The

sensitivity within this region is approximately 0046 pF microm-1 and 0253 pF microm-1 for T1

and T2 respectively The order of magnitude improvement is attributed to the differential

tri-plate arrangement and thicker comb finger

0

02

04

06

08

1

12

14

16

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-4 lsquorsquo represents the output of the differential tri-plate drive while lsquorsquo represents the output of the transverse drive The dashed line represents the linear approximation for displacements less that 1 microm For the purpose of comparison the output of the transverse drive was equalized to zero at zero displacement

21

In this work the transverse drive is first implemented in sensor T1 using the

PolyMUMPS fabrication process Then in an effort to linearize the sensor response the

lateral drive is used in L1 using SoiMUMPS The reduction in drive sensitivity due to the

lateral drive is further reduced by increasing the gap distance to 3 microm which is

recommended by the foundry to ensure a successful device These reductions are

balanced by increasing the number of comb fingers N and selecting a substrate thickness

that is 125 times large than T1 The third device T2 is fabricated using SoiMUMPS and

uses the tri-plate differential drive This design capitalizes on the increase in sensitivity of

a gap closing drive and larger substrate thickness attaining an order of magnitude

increase in sensitivity over T1 and L1 The specifications for each drive along with the

successive improvement in drive sensitivity are outlined in Table 3-1

Table 3-1 Specifications for each capacitive comb drive implemented in this work T1 L1 and T2

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Comb Finger Thickness tc (microm) 2 25 25 Gap Distance x0 (microm) 2 3 2 Number of Fingers N 200 256 30 Overlap length y0 (microm) 45 10 70 Drive Sensitivity (pF microm-1) 0046 0038 0273

32 MECHANICAL ANALYSIS

321 MICRO-FLEXURE SELECTION

The sensor structure consists of a proof mass connected to a compliant micro-flexure

that itself is anchored to the sensor substrate The proof mass displaces in response to an

22

applied acceleration and this displacement is sensed using the capacitive comb drive

attached to the proof mass

Proof mass

Folded beam micro-flexure

Anchor

Figure 3-5 Sensor Structure

This section focuses on micro-flexure selection as it should exhibit a large linear

displacement range and have a high lateral stiffness ratio to reduce cross-axis sensitivity

A number of micro-flexures have been used in MEMS research these include the

clamped-clamped crab-leg and folded beam shown in Figure 3-6-through 3-7

The clamped-clamped flexure exhibits a high stiffness ratio however for large lateral

displacements extensional axial forces buildup within the beams resulting in a non-linear

force-displacement relation [40] Within the linear regime the stiffness of the clamped-

clamped flexure is expressed as [40]

3

3b

yEt wk

L=

(13)

23

where E is the Youngrsquos Modulus of the beam material (N m-2) tb is the beam out of

plane thickness (m) w is the beam width (m) and L is the beam length (m) The stiffness

ratio is expressed as [40]

2x

y

k Lk w

=

(14)

If L is two orders of magnitude greater than w this results in a ratio of approximately

10000 The linear displacement range is determined using Finite Element Analysis (FEA)

of a clamped-clamped flexure with L=387 microm w=4 microm and tb=2 microm The result of this

simulation is plotted in and compared to a structure with constant stiffness ky

= 035 N m-1 predicted by Eq (13) The linear relation is valid for small deflections that

are less than 2 microm In addition this micro-flexure is susceptible to buckling as a result of

a compressive residual stress that may arise due to the fabrication process [41]

Figure 3-9

In the crab leg design the thigh section is meant to relieve the extensional axial

stress postponing the occurrence of non-linearity until greater lateral displacement [42]

The stiffness of the crab leg flexure is expressed as [43]

3

44

b s s ty

s s t

Et w L LkL L L

κκ

+= +

(15)

where κ=IsIt=(wswt)3 is a ratio of the second moment of inertia of the shin (s) and thigh

(t) ws and wt represent the width and Ls and Lt represent the length of the shin and thigh

regions respectively The stiffness ratio is expressed as [43]

34

4t s s tx

y s t s

w L L Lkk w L L L

αα t

+= +

(16)

The linear displacement range of a crab leg flexure with Ls=285 microm Lt=30 microm ws=wt=3

microm and tb=2 microm is determined using FEA The results are plotted in Figure 3-9 and

compared to a structure of constant stiffness ky=035 N m-1 predicted by Eq (15) The

24

non-linearity is greatly reduced when compared to the clamp-clamped flexure with Eq

accurately predicting the deflection until 4 microm However the flexure is more

susceptible to off-axis deflection as there is an order of magnitude reduction in the

stiffness ratio as compared to the clamped-clamped flexure

(15)

The folded beam flexure stands out as it is designed to exhibit a large linear

displacement range and to relieve residual stress due to the manufacturing process by

allowing the beams to expand and contract along the axial direction [40] It was first used

in an electrostatic resonator and has subsequently been used for electrostatic actuation in

a magnetic hard drives and micro-positioning [3344] If the truss section is rigid

ItrussgtgtIbeam then the stiffness of the flexure is expressed as [4043]

3

32b

yEt wk

L=

(17)

The stiffness ratio is expressed as [40]

2x

y

k Lk w

=

(18)

Figure 3-6 Clamped-clamped flexure

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

6

sensitivity gained from a transverse drive but rejects off-axis displacement errors and

common mode using a differential tri-plate arrangement In addition an off-chip

capacitance to voltage readout circuit is fabricated and tested on sensor L1

Chapter 2 begins with an overview on the various methods of fault detection

followed by a review of MEMS vibration sensors each with a different transduction

mechanism Chapter 3 outlines the sensor structure focusing on the mechanical design

and sensing principal Chapter 4 focuses on the readout circuitry layout and simulation

results Chapter 5 outlines the various source of noise including mechanical noise arising

from the sensor and electric noise from the readout circuit Chapter 6 overviews the

foundry fabrication processes used including PolyMUMPs and SoiMUMPS The

experimental results of two sensors T1 and L1 which were fabricated and tested within

the thesis time frame are presented in Chapter 7 Chapter 8 concludes the thesis with a

brief summary and outline for future work

2 LITERATURE REVIEW

21 EXISTING METHODS FOR FAULT DETECTION

As this thesis develops a sensor for fault detection using vibration monitoring this

section overviews three existing methods for fault detection in mechanical systems First

motor current monitoring is discussed as it has the ability to measure significant bearing

faults in motor systems Next temperature sensors specifically thin film thermocouples

(TFTC) and MEMS temperature sensors are discussed These sensors are mounted in the

chamber of a lubricating fluid or in close proximity to a moving component and monitor

temperature fluctuations caused by increased wear Finally acoustic emission (AE)

sensors are discussed which have the ability to detect ultrasonic emissions generated by

plastic deformation in the crystalline lattice of a mechanical component

211 MOTOR CURRENT MONITORING

Monitoring of stator current is used as a method of determining bearing related faults

that produce radial motion between rotor and stator This motion varies the air gap flux

density producing stator currents at predictable frequencies fbng that represent the

bearing faults [10]

bng e vf f m f= plusmn sdot (2)

where fe is the electrical supply frequency m=123hellip and fv is one of the characteristics

frequencies listed in the equations of (1)The bearing damage must be significant to

produce noticeable current fluctuations and it may be difficult to distinguish from current

fluctuations due to voltage imbalances machine saturation or statorarmature faults [2]

7

8

212 TEMPERATURE SENSORS

Temperature sensors are largely used in high temperature high load and high speed

environments typically turbine engines and industrial tools Thermocouple sensors

consist of two dissimilar metals joined together at two junctions and when exposed to a

temperature difference generate a thermoelectric voltage dependant only on the material

properties and junction temperature difference [11] In turbine engines these sensors can

be placed on bearing casings or in the lubricant to monitor temperature fluctuations due

to increased friction The thermocouples are fabricated using bulk materials or micro-

machined thin films (TFTC) with the latter providing a faster response small package

and substantially lower production costs [12] The TFTC can be embedded for in situ

monitoring of processes in hostile environments One study successfully embedded

TFTC into electroplated nickel and attained device characteristics on par with surface

mounted devices [13] More recently MEMS temperature sensors have been developed

that consist of micro-machined semiconductor material that undergoes structural

deformation with temperature changes A common implementation is a bent beam

structure whose temperature induced deflection δ is expressed as [14]

( ) 3

2 2

sin12 cos sinz

EA T LEI AEL

α βδ 2β β

∆=

+

(3)

Temperature sensors are well suited for localized fault monitoring and detection but can

only characterize a mechanical system consisting of multiple components through the use

of multiple sensors

9

213 ACOUSTIC EMISSION SENSORS

Acoustic emission (AE) sensors have been used to characterize wear in machine

tools and monitor bearing and gear problems in centrifugal pumps [1516] First

developed as a Non-Destructive Testing (NDT) technique to detect cracks in civil

structures these sensors detect acoustic emissions generated by the release of vibration

waves in a crystalline lattice due to plastic deformation or crack growth [16]

Measurements are made using piezoelectric transducers with high natural frequencies

100 kHz to 1 MHz to capture the ultrasonic AE emissions An AE sensor is useful as it

has the ability to detect subsurface cracks in gear teeth or bearings before appearing on

the surface causing further damage [17] More recently MEMS acoustic sensors have

been developed and one design includes multiple transducers on a single substrate

which each detect acoustic emission energy at different frequencies [18] This helps

distinguish spurious acoustic emissions arising from impact and friction from those

arising from plastic deformation When compared to typical piezoelectric sensors the

MEMS devices have lower sensitivities and fail to detect some acoustic emissions

[1819] In addition the acoustic emission signal suffers severe attenuation as is crosses

various interfaces such as a gearbox or bearing casings In one experiment consisting of

a pinion gear and an associated bearing a 44dB attenuation was seen between an AE

sensor placed directly on the pinion to one placed on the bearing casing and in some

cases intermediate loss of the signal was observed [20]

10

22 MEMS VIBRATION SENSORS

This thesis focuses on the development of a MEMS vibration sensor for fault

detection as the vibration signature of a mechanical system has the potential to give an

overview of the entire system Many types of MEMS vibration sensors exist with

different operating principals based on resonance the piezoelectric effect and

displacement variation

221 RESONANT SENSORS

Resonant sensors are based on the resonant frequency shift of a fixed-fixed micro-

beam that is excited into resonance using electrostatic thermal or piezoelectric methods

Under an applied axial load there is a shift in the resonant frequency which is expressed

as [21]

2 2

2

1 12 1

nn n

EI Fl2

fl A Eα γ

π ρ= +

I

(4)

where l is the beam length E is Youngrsquos Modulus I is the moment of inertia ρ is the

density A is the cross section area F is the applied axial force αn and γn are mode

dependant coefficients and are 4730 and 0295 for mode 1 in the case discussed The

axial force F is generated by a proof mass attached to one beam end which moves in

response to an external acceleration The performance of resonant sensors are dependant

on the quality factor Q which is the ratio of total energy stored in the system to the

energy lost per cycle and determines the sharpness and amplitude of the resonance peak

A high Q resonator develops a peak that is easily distinguishable in a phase locked loop

control circuit and allows for improved performance and resolution [3] A resonator Q

value is lowered by energy losses to the surrounding fluid attached supports and

11

material Losses to the fluid surrounding the MEMS structure is the dominant loss

mechanism for MEMS sensors operated in air and can be greatly reduced by vacuum

packaging which can be accomplished using glass-silicon anodic bonding [22] In

addition to increasing fabrication cost the long term stability of the vacuum degrades

with ultimate device failure stemming from leakage through micro cracks and defects

and out-gassing [22] Support losses arise from the restoring forces they generate and

can be reduce by balanced designs such as a double ended tuning fork (DETF) or

operation in higher modes [3] In addition the sensitivity of a resonant sensor can be

increased by the addition of micro-levers to increase the axial force In one such design a

two stage micro-lever increased the sensitivity by an order of magnitude over a more

conventional single lever arm [2324]

222 PIEZOELECTRIC SENSORS

Piezoelectric materials produce a charge in response to an applied force This is an

intrinsic effect in materials such as Zinc Oxide (ZnO) whereas materials such as Barium

Titanate or Lead Zirconate Titanate (PZT) need to be poled by placing the material in a

strong electric field at an elevated temperature [3] The piezoelectric effect is expressed

mathematical by [25]

i ij j ik kD d T Eε= + (5)

where D represents the electrical displacement (Cm2) d is the charge coefficient (CN)

T is the applied stress matrix (Nm2) ε is the permittivity matrix (Fm) and E is the

electric field matrix (Vm) The subscripts i=123 j=123456 and k=123 represent

the direction to which the physical property is related For micro-sensors poled in the out

of plane direction only the D3 term is relevant and if fabricated using ZnO or PZT only

12

the charge coefficients d31 d32 and d33 are relevant [25] The piezoelectric material is

commonly deposited as a thin film on the surface of compliant structures in bulk micro-

machined sensors or incorporated as a layer in surface micro-machining on a similar

structure [26] However the inclusion of multiple materials introduces large temperature

variations causing out of plane bending especially in surface micro-machined devices

[2627] Poled piezoelectric materials could be depolarized which serves to reduce the

piezoelectric affect in the material This can occur if the sensor is exposed to a strong

electric field of the opposite polarity high temperatures in excess of the Curie point or

high mechanical stress Also the piezoelectric effect reduces over time an effect that can

be reduced with the addition of composite elements [28] Foundry processes such as

PolyMUMPS and SoiMUMPS do not incorporate piezoelectric layers thus sensors using

this method of transduction must use a customized process The piezoelectric sensor

outlined in [26] uses polysilicon surface micro-machining similar to PolyMUMPS

however it incorporates a ZnO layer that is not available in the foundry process

223 DISPLACEMENT VARIATION SENSORS

Displacement variation sensors consist of a proof mass connected to a compliant

micro-flexure The movement of the proof mass is sensed using optical electron

tunneling or more commonly capacitive methods A recent optical accelerometer

utilized a novel multilayer nano-grating for in-plane displacement sensing of the proof

mass [29] For small space variations between the nano-gratings a large change in the

optical reflection amplitude of the grating was observed [30] Even though optical

sensors offer resolutions approaching the Brownian noise limit they are not available in

small packages that are easily placed in space constricted areas for a low cost [29]

13

The electron tunneling effect is observed when the proximity between two electrodes

one flexible and the other fixed is on the order of 10Ǻ [31] If the two electrodes are

biased with voltage VB a tunneling current It flows between the gap [31]

exp( )t B I tgI V xαinfin sdot minus Φ (6)

where α1 is a constant and equal to 1025 Ǻ-1 eV-frac12 Ф is the effective height of the

tunneling barrier and xtg is the tunneling gap When the flexible electrode moves in

response to acceleration a feedback controller maintains the original distance while

determining the applied acceleration Electron tunneling sensors are able to sense

accelerations in the microg-ng range due to the exponential relation between tunneling current

and gap However these highly sensitive devices are susceptible to thermo-mechanical

noise which could be reduced by operation at low pressure [32] It is also important for

the two electrodes to be coated with metal which is difficult to achieve with foundry

fabrication

Capacitive sensors are based on the parallel plate capacitor and are implemented

using inter-digitated comb fingers in an in-plane lateral or transverse drive or in an out of

plane parallel plate drive [33] These sensors have the advantage of high sensitivity low

power consumption small package low temperature dependence and are easily

integrated with existing foundry fabrication processes [34] The lateral drive offers a fully

linear response but low sensitivity whereas the transverse and parallel plate drives

achieve high sensitivity if gap distances are of a few microns and large capacitance areas

are utilized However both drives have a non-linear response and are prone to pull-in

Also if these sensors are fabricated using a surface micromachining process such as

PolyMUMPS where the ground plane is a few microns above the structural layer they

14

are prone to electrical failure due to shorting This was observed in an experiment where

similar sensors were placed in a vibrating environment with a peak of 120g and shorting

between the structural layer and ground plane was found to be the predominate mode of

failure [35] However with certain fabrication processes such as SoiMUMPS this issue

can be neglected as the substrate material below any moving component is removed

Most capacitance sensors are fabricated using an integrated CMOS (Complementary

Metal Oxide Semiconductor)-MEMS technique where the MEMS sensor is fabricated

before during or after the CMOS circuit fabrication [36] Although these sensors offer

high-sensitivity and small parasitic capacitance they suffer from in-plane and out of

plane curling of the beam which reduces the capacitance between adjacent comb fingers

[37] In addition multiple processing steps are required for device fabrication

More than the other transduction methods MEMS capacitive displacement sensors are

attractive for practical implementation in an automobile and therefore is the selected

method for the vibration sensor presented in this work Their low temperature dependence

is an ideal characteristic for an automobile environment that faces varying temperature If

implemented with off-chip readout circuitry they require a standard foundry fabrication

process which provides a strong platform for mass production and low unit cost

3 SENSOR STRUCTURE

This chapter presents the physical structure of the three MEMS vibration sensors

which includes the displacement sensing capacitive comb drive and the micro-

flexureproof mass system First the capacitive comb drive and the three variations

implemented the transverse drive lateral drive and differential tri-plate drive are

discussed Next the mechanical structure is presented and includes the micro-flexure

selection quality factor and equations of motion Finally instability that places a limit

on the operating range of the sensor as a result of a gap closing capacitive comb drive is

discussed

31 SENSING

Capacitive sensors are based on the parallel plate capacitor and are commonly

implemented using inter-digitated comb fingers A capacitance can be realized if a voltage

is placed between two closely spaced plates If the fringing electric field is neglected the

capacitance of the parallel plate capacitor Cpp (F) is expressed as [28]

air pppp

k AC

dε

= (7)

where kair=1 is the dielectric constant of air at atmospheric pressure ε=88542times10-12 F m-1

is the permittivity of free space App is the overlapping area (m2) and d is the distance

between the two plates (m)

15

16

Figure 3-1 An inter-digitated comb drive where changes in capacitance are generated by either changes in gap distance x0 or in the overlap area y0timest

Figure 3-1

For a set of inter-digitated comb fingers or comb drive shown in the total

capacitance Ctotal is expressed as [28]

0 0

0 0

( ) ( )( ) (Total cy y y yC N t

)x x x xε

+ += + + minus

(8)

where N is the number of comb fingers on the movable comb y0 is the initial comb finger

overlap length (m) tc is the out of plane comb finger thickness (m) and x0 is the inital

finger gap (m) In this work the comb drive is implemented in three arrangements

transverse lateral and differential tri-plate In a transverse comb drive changes in

capacitance are generated by transverse movements along the x-axis resulting in changes

to the gap x0 between the movable and fixed comb fingers The nonlinear relation

between total capacitance CTotalT and displacement is expressed as

17

00 0

1 1( ) (Total TC N ty

)x x x xε

= + minus +

(9)

Although the drive sensitivity part is not constant an average sensitivity can be

assumed for small displacements [38]

Total TC xpart

In a lateral comb drive changes in capacitance are generated by lateral movements

along the y-axis resulting in changes to the overlapping area y0timestc The total capacitance

CTotalL between all comb fingers is expressed as

00

2 ( )Total LN tC yx

yε= +

(10)

The sensitivity of the lateral comb drive SL is expressed as

0

2Total LL

C N tSy x

εpart= =

part

(11)

The lateral comb drive was first introduced by W C Tang T H Nguyen and R T

Howe [33] in an electrostatic resonator and has since been used numerous times in both

drive and sense configurations It has the advantage of a fully linear response and

negligible instability but offers low sensitivity The sensitivity for the lateral drive L1

implemented in this work is 0038 pF microm-1

A differential tri-plate drive shown in Figure 3-2 which is a variation of the

transverse drive is implemented to improve sensitivity and reduce noise [39]

18

Figure 3-2 An inter-digitated comb drive in a differential tri-plate arrangement where x01 x02 and x03 are the gap distances and x01= x02ltlt x03 to improve sensitivity

In the tri-plate drive the at rest gap distances between the two capacitors formed by any

movable finger is not equivalent For a differential arrangement insuring that x01=

x02ltltx03 the linearity for small displacements (less than 1 microm) is enhanced over the case

of a single ended arrangement where x01=x02=x03=x0 which is equivalent to the

transverse drive This is shown graphically in Figure 3-3 where x01= x02=x03=2 microm in

the lower curve and x01= x02= 2 microm and x03= 15 microm in the upper curve A substantial

improvement in linearity is attained with the tri-plate drive

19

0

005

01

015

02

025

03

035

0 02 04 06 08 1

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-3 Comparison of the effect of gap distances on linearity for two case lsquorsquo represents case 1 where the fingers in are implemented in a differential tri-plate arrangement and x01= x02= 2 microm and x03= 15 microm lsquorsquo represents the case 2 where the fingers are implemented in a single ended arrangement and x01= x02=x03=2 microm

The tri-plate drive is also implemented in a differential arrangement where the output

is the change in capacitance rather than the total capacitance resulting from an external

excitation This reduces errors cause by off-axis excitation and common mode noise The

output of the differential tri-plate drive is expressed as [39]

Total Top Total BottomC C C∆ = minus where

003 01

1 1( ) (Total TopC N ty

)x x x xε

= +

minus +

(12)

20

002 03

1 1( ) (Total BottomC N ty

)x x x xε

= +

minus +

A graphic comparison of the capacitancedisplacement relation for sensors T1 and T2 is

shown in Figure 3-4 For these gap closing drives it is assumed that for displacements

less that 1 microm the relation is linear indicated by the dashed line above each curve The

sensitivity within this region is approximately 0046 pF microm-1 and 0253 pF microm-1 for T1

and T2 respectively The order of magnitude improvement is attributed to the differential

tri-plate arrangement and thicker comb finger

0

02

04

06

08

1

12

14

16

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-4 lsquorsquo represents the output of the differential tri-plate drive while lsquorsquo represents the output of the transverse drive The dashed line represents the linear approximation for displacements less that 1 microm For the purpose of comparison the output of the transverse drive was equalized to zero at zero displacement

21

In this work the transverse drive is first implemented in sensor T1 using the

PolyMUMPS fabrication process Then in an effort to linearize the sensor response the

lateral drive is used in L1 using SoiMUMPS The reduction in drive sensitivity due to the

lateral drive is further reduced by increasing the gap distance to 3 microm which is

recommended by the foundry to ensure a successful device These reductions are

balanced by increasing the number of comb fingers N and selecting a substrate thickness

that is 125 times large than T1 The third device T2 is fabricated using SoiMUMPS and

uses the tri-plate differential drive This design capitalizes on the increase in sensitivity of

a gap closing drive and larger substrate thickness attaining an order of magnitude

increase in sensitivity over T1 and L1 The specifications for each drive along with the

successive improvement in drive sensitivity are outlined in Table 3-1

Table 3-1 Specifications for each capacitive comb drive implemented in this work T1 L1 and T2

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Comb Finger Thickness tc (microm) 2 25 25 Gap Distance x0 (microm) 2 3 2 Number of Fingers N 200 256 30 Overlap length y0 (microm) 45 10 70 Drive Sensitivity (pF microm-1) 0046 0038 0273

32 MECHANICAL ANALYSIS

321 MICRO-FLEXURE SELECTION

The sensor structure consists of a proof mass connected to a compliant micro-flexure

that itself is anchored to the sensor substrate The proof mass displaces in response to an

22

applied acceleration and this displacement is sensed using the capacitive comb drive

attached to the proof mass

Proof mass

Folded beam micro-flexure

Anchor

Figure 3-5 Sensor Structure

This section focuses on micro-flexure selection as it should exhibit a large linear

displacement range and have a high lateral stiffness ratio to reduce cross-axis sensitivity

A number of micro-flexures have been used in MEMS research these include the

clamped-clamped crab-leg and folded beam shown in Figure 3-6-through 3-7

The clamped-clamped flexure exhibits a high stiffness ratio however for large lateral

displacements extensional axial forces buildup within the beams resulting in a non-linear

force-displacement relation [40] Within the linear regime the stiffness of the clamped-

clamped flexure is expressed as [40]

3

3b

yEt wk

L=

(13)

23

where E is the Youngrsquos Modulus of the beam material (N m-2) tb is the beam out of

plane thickness (m) w is the beam width (m) and L is the beam length (m) The stiffness

ratio is expressed as [40]

2x

y

k Lk w

=

(14)

If L is two orders of magnitude greater than w this results in a ratio of approximately

10000 The linear displacement range is determined using Finite Element Analysis (FEA)

of a clamped-clamped flexure with L=387 microm w=4 microm and tb=2 microm The result of this

simulation is plotted in and compared to a structure with constant stiffness ky

= 035 N m-1 predicted by Eq (13) The linear relation is valid for small deflections that

are less than 2 microm In addition this micro-flexure is susceptible to buckling as a result of

a compressive residual stress that may arise due to the fabrication process [41]

Figure 3-9

In the crab leg design the thigh section is meant to relieve the extensional axial

stress postponing the occurrence of non-linearity until greater lateral displacement [42]

The stiffness of the crab leg flexure is expressed as [43]

3

44

b s s ty

s s t

Et w L LkL L L

κκ

+= +

(15)

where κ=IsIt=(wswt)3 is a ratio of the second moment of inertia of the shin (s) and thigh

(t) ws and wt represent the width and Ls and Lt represent the length of the shin and thigh

regions respectively The stiffness ratio is expressed as [43]

34

4t s s tx

y s t s

w L L Lkk w L L L

αα t

+= +

(16)

The linear displacement range of a crab leg flexure with Ls=285 microm Lt=30 microm ws=wt=3

microm and tb=2 microm is determined using FEA The results are plotted in Figure 3-9 and

compared to a structure of constant stiffness ky=035 N m-1 predicted by Eq (15) The

24

non-linearity is greatly reduced when compared to the clamp-clamped flexure with Eq

accurately predicting the deflection until 4 microm However the flexure is more

susceptible to off-axis deflection as there is an order of magnitude reduction in the

stiffness ratio as compared to the clamped-clamped flexure

(15)

The folded beam flexure stands out as it is designed to exhibit a large linear

displacement range and to relieve residual stress due to the manufacturing process by

allowing the beams to expand and contract along the axial direction [40] It was first used

in an electrostatic resonator and has subsequently been used for electrostatic actuation in

a magnetic hard drives and micro-positioning [3344] If the truss section is rigid

ItrussgtgtIbeam then the stiffness of the flexure is expressed as [4043]

3

32b

yEt wk

L=

(17)

The stiffness ratio is expressed as [40]

2x

y

k Lk w

=

(18)

Figure 3-6 Clamped-clamped flexure

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

2 LITERATURE REVIEW

21 EXISTING METHODS FOR FAULT DETECTION

As this thesis develops a sensor for fault detection using vibration monitoring this

section overviews three existing methods for fault detection in mechanical systems First

motor current monitoring is discussed as it has the ability to measure significant bearing

faults in motor systems Next temperature sensors specifically thin film thermocouples

(TFTC) and MEMS temperature sensors are discussed These sensors are mounted in the

chamber of a lubricating fluid or in close proximity to a moving component and monitor

temperature fluctuations caused by increased wear Finally acoustic emission (AE)

sensors are discussed which have the ability to detect ultrasonic emissions generated by

plastic deformation in the crystalline lattice of a mechanical component

211 MOTOR CURRENT MONITORING

Monitoring of stator current is used as a method of determining bearing related faults

that produce radial motion between rotor and stator This motion varies the air gap flux

density producing stator currents at predictable frequencies fbng that represent the

bearing faults [10]

bng e vf f m f= plusmn sdot (2)

where fe is the electrical supply frequency m=123hellip and fv is one of the characteristics

frequencies listed in the equations of (1)The bearing damage must be significant to

produce noticeable current fluctuations and it may be difficult to distinguish from current

fluctuations due to voltage imbalances machine saturation or statorarmature faults [2]

7

8

212 TEMPERATURE SENSORS

Temperature sensors are largely used in high temperature high load and high speed

environments typically turbine engines and industrial tools Thermocouple sensors

consist of two dissimilar metals joined together at two junctions and when exposed to a

temperature difference generate a thermoelectric voltage dependant only on the material

properties and junction temperature difference [11] In turbine engines these sensors can

be placed on bearing casings or in the lubricant to monitor temperature fluctuations due

to increased friction The thermocouples are fabricated using bulk materials or micro-

machined thin films (TFTC) with the latter providing a faster response small package

and substantially lower production costs [12] The TFTC can be embedded for in situ

monitoring of processes in hostile environments One study successfully embedded

TFTC into electroplated nickel and attained device characteristics on par with surface

mounted devices [13] More recently MEMS temperature sensors have been developed

that consist of micro-machined semiconductor material that undergoes structural

deformation with temperature changes A common implementation is a bent beam

structure whose temperature induced deflection δ is expressed as [14]

( ) 3

2 2

sin12 cos sinz

EA T LEI AEL

α βδ 2β β

∆=

+

(3)

Temperature sensors are well suited for localized fault monitoring and detection but can

only characterize a mechanical system consisting of multiple components through the use

of multiple sensors

9

213 ACOUSTIC EMISSION SENSORS

Acoustic emission (AE) sensors have been used to characterize wear in machine

tools and monitor bearing and gear problems in centrifugal pumps [1516] First

developed as a Non-Destructive Testing (NDT) technique to detect cracks in civil

structures these sensors detect acoustic emissions generated by the release of vibration

waves in a crystalline lattice due to plastic deformation or crack growth [16]

Measurements are made using piezoelectric transducers with high natural frequencies

100 kHz to 1 MHz to capture the ultrasonic AE emissions An AE sensor is useful as it

has the ability to detect subsurface cracks in gear teeth or bearings before appearing on

the surface causing further damage [17] More recently MEMS acoustic sensors have

been developed and one design includes multiple transducers on a single substrate

which each detect acoustic emission energy at different frequencies [18] This helps

distinguish spurious acoustic emissions arising from impact and friction from those

arising from plastic deformation When compared to typical piezoelectric sensors the

MEMS devices have lower sensitivities and fail to detect some acoustic emissions

[1819] In addition the acoustic emission signal suffers severe attenuation as is crosses

various interfaces such as a gearbox or bearing casings In one experiment consisting of

a pinion gear and an associated bearing a 44dB attenuation was seen between an AE

sensor placed directly on the pinion to one placed on the bearing casing and in some

cases intermediate loss of the signal was observed [20]

10

22 MEMS VIBRATION SENSORS

This thesis focuses on the development of a MEMS vibration sensor for fault

detection as the vibration signature of a mechanical system has the potential to give an

overview of the entire system Many types of MEMS vibration sensors exist with

different operating principals based on resonance the piezoelectric effect and

displacement variation

221 RESONANT SENSORS

Resonant sensors are based on the resonant frequency shift of a fixed-fixed micro-

beam that is excited into resonance using electrostatic thermal or piezoelectric methods

Under an applied axial load there is a shift in the resonant frequency which is expressed

as [21]

2 2

2

1 12 1

nn n

EI Fl2

fl A Eα γ

π ρ= +

I

(4)

where l is the beam length E is Youngrsquos Modulus I is the moment of inertia ρ is the

density A is the cross section area F is the applied axial force αn and γn are mode

dependant coefficients and are 4730 and 0295 for mode 1 in the case discussed The

axial force F is generated by a proof mass attached to one beam end which moves in

response to an external acceleration The performance of resonant sensors are dependant

on the quality factor Q which is the ratio of total energy stored in the system to the

energy lost per cycle and determines the sharpness and amplitude of the resonance peak

A high Q resonator develops a peak that is easily distinguishable in a phase locked loop

control circuit and allows for improved performance and resolution [3] A resonator Q

value is lowered by energy losses to the surrounding fluid attached supports and

11

material Losses to the fluid surrounding the MEMS structure is the dominant loss

mechanism for MEMS sensors operated in air and can be greatly reduced by vacuum

packaging which can be accomplished using glass-silicon anodic bonding [22] In

addition to increasing fabrication cost the long term stability of the vacuum degrades

with ultimate device failure stemming from leakage through micro cracks and defects

and out-gassing [22] Support losses arise from the restoring forces they generate and

can be reduce by balanced designs such as a double ended tuning fork (DETF) or

operation in higher modes [3] In addition the sensitivity of a resonant sensor can be

increased by the addition of micro-levers to increase the axial force In one such design a

two stage micro-lever increased the sensitivity by an order of magnitude over a more

conventional single lever arm [2324]

222 PIEZOELECTRIC SENSORS

Piezoelectric materials produce a charge in response to an applied force This is an

intrinsic effect in materials such as Zinc Oxide (ZnO) whereas materials such as Barium

Titanate or Lead Zirconate Titanate (PZT) need to be poled by placing the material in a

strong electric field at an elevated temperature [3] The piezoelectric effect is expressed

mathematical by [25]

i ij j ik kD d T Eε= + (5)

where D represents the electrical displacement (Cm2) d is the charge coefficient (CN)

T is the applied stress matrix (Nm2) ε is the permittivity matrix (Fm) and E is the

electric field matrix (Vm) The subscripts i=123 j=123456 and k=123 represent

the direction to which the physical property is related For micro-sensors poled in the out

of plane direction only the D3 term is relevant and if fabricated using ZnO or PZT only

12

the charge coefficients d31 d32 and d33 are relevant [25] The piezoelectric material is

commonly deposited as a thin film on the surface of compliant structures in bulk micro-

machined sensors or incorporated as a layer in surface micro-machining on a similar

structure [26] However the inclusion of multiple materials introduces large temperature

variations causing out of plane bending especially in surface micro-machined devices

[2627] Poled piezoelectric materials could be depolarized which serves to reduce the

piezoelectric affect in the material This can occur if the sensor is exposed to a strong

electric field of the opposite polarity high temperatures in excess of the Curie point or

high mechanical stress Also the piezoelectric effect reduces over time an effect that can

be reduced with the addition of composite elements [28] Foundry processes such as

PolyMUMPS and SoiMUMPS do not incorporate piezoelectric layers thus sensors using

this method of transduction must use a customized process The piezoelectric sensor

outlined in [26] uses polysilicon surface micro-machining similar to PolyMUMPS

however it incorporates a ZnO layer that is not available in the foundry process

223 DISPLACEMENT VARIATION SENSORS

Displacement variation sensors consist of a proof mass connected to a compliant

micro-flexure The movement of the proof mass is sensed using optical electron

tunneling or more commonly capacitive methods A recent optical accelerometer

utilized a novel multilayer nano-grating for in-plane displacement sensing of the proof

mass [29] For small space variations between the nano-gratings a large change in the

optical reflection amplitude of the grating was observed [30] Even though optical

sensors offer resolutions approaching the Brownian noise limit they are not available in

small packages that are easily placed in space constricted areas for a low cost [29]

13

The electron tunneling effect is observed when the proximity between two electrodes

one flexible and the other fixed is on the order of 10Ǻ [31] If the two electrodes are

biased with voltage VB a tunneling current It flows between the gap [31]

exp( )t B I tgI V xαinfin sdot minus Φ (6)

where α1 is a constant and equal to 1025 Ǻ-1 eV-frac12 Ф is the effective height of the

tunneling barrier and xtg is the tunneling gap When the flexible electrode moves in

response to acceleration a feedback controller maintains the original distance while

determining the applied acceleration Electron tunneling sensors are able to sense

accelerations in the microg-ng range due to the exponential relation between tunneling current

and gap However these highly sensitive devices are susceptible to thermo-mechanical

noise which could be reduced by operation at low pressure [32] It is also important for

the two electrodes to be coated with metal which is difficult to achieve with foundry

fabrication

Capacitive sensors are based on the parallel plate capacitor and are implemented

using inter-digitated comb fingers in an in-plane lateral or transverse drive or in an out of

plane parallel plate drive [33] These sensors have the advantage of high sensitivity low

power consumption small package low temperature dependence and are easily

integrated with existing foundry fabrication processes [34] The lateral drive offers a fully

linear response but low sensitivity whereas the transverse and parallel plate drives

achieve high sensitivity if gap distances are of a few microns and large capacitance areas

are utilized However both drives have a non-linear response and are prone to pull-in

Also if these sensors are fabricated using a surface micromachining process such as

PolyMUMPS where the ground plane is a few microns above the structural layer they

14

are prone to electrical failure due to shorting This was observed in an experiment where

similar sensors were placed in a vibrating environment with a peak of 120g and shorting

between the structural layer and ground plane was found to be the predominate mode of

failure [35] However with certain fabrication processes such as SoiMUMPS this issue

can be neglected as the substrate material below any moving component is removed

Most capacitance sensors are fabricated using an integrated CMOS (Complementary

Metal Oxide Semiconductor)-MEMS technique where the MEMS sensor is fabricated

before during or after the CMOS circuit fabrication [36] Although these sensors offer

high-sensitivity and small parasitic capacitance they suffer from in-plane and out of

plane curling of the beam which reduces the capacitance between adjacent comb fingers

[37] In addition multiple processing steps are required for device fabrication

More than the other transduction methods MEMS capacitive displacement sensors are

attractive for practical implementation in an automobile and therefore is the selected

method for the vibration sensor presented in this work Their low temperature dependence

is an ideal characteristic for an automobile environment that faces varying temperature If

implemented with off-chip readout circuitry they require a standard foundry fabrication

process which provides a strong platform for mass production and low unit cost

3 SENSOR STRUCTURE

This chapter presents the physical structure of the three MEMS vibration sensors

which includes the displacement sensing capacitive comb drive and the micro-

flexureproof mass system First the capacitive comb drive and the three variations

implemented the transverse drive lateral drive and differential tri-plate drive are

discussed Next the mechanical structure is presented and includes the micro-flexure

selection quality factor and equations of motion Finally instability that places a limit

on the operating range of the sensor as a result of a gap closing capacitive comb drive is

discussed

31 SENSING

Capacitive sensors are based on the parallel plate capacitor and are commonly

implemented using inter-digitated comb fingers A capacitance can be realized if a voltage

is placed between two closely spaced plates If the fringing electric field is neglected the

capacitance of the parallel plate capacitor Cpp (F) is expressed as [28]

air pppp

k AC

dε

= (7)

where kair=1 is the dielectric constant of air at atmospheric pressure ε=88542times10-12 F m-1

is the permittivity of free space App is the overlapping area (m2) and d is the distance

between the two plates (m)

15

16

Figure 3-1 An inter-digitated comb drive where changes in capacitance are generated by either changes in gap distance x0 or in the overlap area y0timest

Figure 3-1

For a set of inter-digitated comb fingers or comb drive shown in the total

capacitance Ctotal is expressed as [28]

0 0

0 0

( ) ( )( ) (Total cy y y yC N t

)x x x xε

+ += + + minus

(8)

where N is the number of comb fingers on the movable comb y0 is the initial comb finger

overlap length (m) tc is the out of plane comb finger thickness (m) and x0 is the inital

finger gap (m) In this work the comb drive is implemented in three arrangements

transverse lateral and differential tri-plate In a transverse comb drive changes in

capacitance are generated by transverse movements along the x-axis resulting in changes

to the gap x0 between the movable and fixed comb fingers The nonlinear relation

between total capacitance CTotalT and displacement is expressed as

17

00 0

1 1( ) (Total TC N ty

)x x x xε

= + minus +

(9)

Although the drive sensitivity part is not constant an average sensitivity can be

assumed for small displacements [38]

Total TC xpart

In a lateral comb drive changes in capacitance are generated by lateral movements

along the y-axis resulting in changes to the overlapping area y0timestc The total capacitance

CTotalL between all comb fingers is expressed as

00

2 ( )Total LN tC yx

yε= +

(10)

The sensitivity of the lateral comb drive SL is expressed as

0

2Total LL

C N tSy x

εpart= =

part

(11)

The lateral comb drive was first introduced by W C Tang T H Nguyen and R T

Howe [33] in an electrostatic resonator and has since been used numerous times in both

drive and sense configurations It has the advantage of a fully linear response and

negligible instability but offers low sensitivity The sensitivity for the lateral drive L1

implemented in this work is 0038 pF microm-1

A differential tri-plate drive shown in Figure 3-2 which is a variation of the

transverse drive is implemented to improve sensitivity and reduce noise [39]

18

Figure 3-2 An inter-digitated comb drive in a differential tri-plate arrangement where x01 x02 and x03 are the gap distances and x01= x02ltlt x03 to improve sensitivity

In the tri-plate drive the at rest gap distances between the two capacitors formed by any

movable finger is not equivalent For a differential arrangement insuring that x01=

x02ltltx03 the linearity for small displacements (less than 1 microm) is enhanced over the case

of a single ended arrangement where x01=x02=x03=x0 which is equivalent to the

transverse drive This is shown graphically in Figure 3-3 where x01= x02=x03=2 microm in

the lower curve and x01= x02= 2 microm and x03= 15 microm in the upper curve A substantial

improvement in linearity is attained with the tri-plate drive

19

0

005

01

015

02

025

03

035

0 02 04 06 08 1

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-3 Comparison of the effect of gap distances on linearity for two case lsquorsquo represents case 1 where the fingers in are implemented in a differential tri-plate arrangement and x01= x02= 2 microm and x03= 15 microm lsquorsquo represents the case 2 where the fingers are implemented in a single ended arrangement and x01= x02=x03=2 microm

The tri-plate drive is also implemented in a differential arrangement where the output

is the change in capacitance rather than the total capacitance resulting from an external

excitation This reduces errors cause by off-axis excitation and common mode noise The

output of the differential tri-plate drive is expressed as [39]

Total Top Total BottomC C C∆ = minus where

003 01

1 1( ) (Total TopC N ty

)x x x xε

= +

minus +

(12)

20

002 03

1 1( ) (Total BottomC N ty

)x x x xε

= +

minus +

A graphic comparison of the capacitancedisplacement relation for sensors T1 and T2 is

shown in Figure 3-4 For these gap closing drives it is assumed that for displacements

less that 1 microm the relation is linear indicated by the dashed line above each curve The

sensitivity within this region is approximately 0046 pF microm-1 and 0253 pF microm-1 for T1

and T2 respectively The order of magnitude improvement is attributed to the differential

tri-plate arrangement and thicker comb finger

0

02

04

06

08

1

12

14

16

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-4 lsquorsquo represents the output of the differential tri-plate drive while lsquorsquo represents the output of the transverse drive The dashed line represents the linear approximation for displacements less that 1 microm For the purpose of comparison the output of the transverse drive was equalized to zero at zero displacement

21

In this work the transverse drive is first implemented in sensor T1 using the

PolyMUMPS fabrication process Then in an effort to linearize the sensor response the

lateral drive is used in L1 using SoiMUMPS The reduction in drive sensitivity due to the

lateral drive is further reduced by increasing the gap distance to 3 microm which is

recommended by the foundry to ensure a successful device These reductions are

balanced by increasing the number of comb fingers N and selecting a substrate thickness

that is 125 times large than T1 The third device T2 is fabricated using SoiMUMPS and

uses the tri-plate differential drive This design capitalizes on the increase in sensitivity of

a gap closing drive and larger substrate thickness attaining an order of magnitude

increase in sensitivity over T1 and L1 The specifications for each drive along with the

successive improvement in drive sensitivity are outlined in Table 3-1

Table 3-1 Specifications for each capacitive comb drive implemented in this work T1 L1 and T2

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Comb Finger Thickness tc (microm) 2 25 25 Gap Distance x0 (microm) 2 3 2 Number of Fingers N 200 256 30 Overlap length y0 (microm) 45 10 70 Drive Sensitivity (pF microm-1) 0046 0038 0273

32 MECHANICAL ANALYSIS

321 MICRO-FLEXURE SELECTION

The sensor structure consists of a proof mass connected to a compliant micro-flexure

that itself is anchored to the sensor substrate The proof mass displaces in response to an

22

applied acceleration and this displacement is sensed using the capacitive comb drive

attached to the proof mass

Proof mass

Folded beam micro-flexure

Anchor

Figure 3-5 Sensor Structure

This section focuses on micro-flexure selection as it should exhibit a large linear

displacement range and have a high lateral stiffness ratio to reduce cross-axis sensitivity

A number of micro-flexures have been used in MEMS research these include the

clamped-clamped crab-leg and folded beam shown in Figure 3-6-through 3-7

The clamped-clamped flexure exhibits a high stiffness ratio however for large lateral

displacements extensional axial forces buildup within the beams resulting in a non-linear

force-displacement relation [40] Within the linear regime the stiffness of the clamped-

clamped flexure is expressed as [40]

3

3b

yEt wk

L=

(13)

23

where E is the Youngrsquos Modulus of the beam material (N m-2) tb is the beam out of

plane thickness (m) w is the beam width (m) and L is the beam length (m) The stiffness

ratio is expressed as [40]

2x

y

k Lk w

=

(14)

If L is two orders of magnitude greater than w this results in a ratio of approximately

10000 The linear displacement range is determined using Finite Element Analysis (FEA)

of a clamped-clamped flexure with L=387 microm w=4 microm and tb=2 microm The result of this

simulation is plotted in and compared to a structure with constant stiffness ky

= 035 N m-1 predicted by Eq (13) The linear relation is valid for small deflections that

are less than 2 microm In addition this micro-flexure is susceptible to buckling as a result of

a compressive residual stress that may arise due to the fabrication process [41]

Figure 3-9

In the crab leg design the thigh section is meant to relieve the extensional axial

stress postponing the occurrence of non-linearity until greater lateral displacement [42]

The stiffness of the crab leg flexure is expressed as [43]

3

44

b s s ty

s s t

Et w L LkL L L

κκ

+= +

(15)

where κ=IsIt=(wswt)3 is a ratio of the second moment of inertia of the shin (s) and thigh

(t) ws and wt represent the width and Ls and Lt represent the length of the shin and thigh

regions respectively The stiffness ratio is expressed as [43]

34

4t s s tx

y s t s

w L L Lkk w L L L

αα t

+= +

(16)

The linear displacement range of a crab leg flexure with Ls=285 microm Lt=30 microm ws=wt=3

microm and tb=2 microm is determined using FEA The results are plotted in Figure 3-9 and

compared to a structure of constant stiffness ky=035 N m-1 predicted by Eq (15) The

24

non-linearity is greatly reduced when compared to the clamp-clamped flexure with Eq

accurately predicting the deflection until 4 microm However the flexure is more

susceptible to off-axis deflection as there is an order of magnitude reduction in the

stiffness ratio as compared to the clamped-clamped flexure

(15)

The folded beam flexure stands out as it is designed to exhibit a large linear

displacement range and to relieve residual stress due to the manufacturing process by

allowing the beams to expand and contract along the axial direction [40] It was first used

in an electrostatic resonator and has subsequently been used for electrostatic actuation in

a magnetic hard drives and micro-positioning [3344] If the truss section is rigid

ItrussgtgtIbeam then the stiffness of the flexure is expressed as [4043]

3

32b

yEt wk

L=

(17)

The stiffness ratio is expressed as [40]

2x

y

k Lk w

=

(18)

Figure 3-6 Clamped-clamped flexure

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

8

212 TEMPERATURE SENSORS

Temperature sensors are largely used in high temperature high load and high speed

environments typically turbine engines and industrial tools Thermocouple sensors

consist of two dissimilar metals joined together at two junctions and when exposed to a

temperature difference generate a thermoelectric voltage dependant only on the material

properties and junction temperature difference [11] In turbine engines these sensors can

be placed on bearing casings or in the lubricant to monitor temperature fluctuations due

to increased friction The thermocouples are fabricated using bulk materials or micro-

machined thin films (TFTC) with the latter providing a faster response small package

and substantially lower production costs [12] The TFTC can be embedded for in situ

monitoring of processes in hostile environments One study successfully embedded

TFTC into electroplated nickel and attained device characteristics on par with surface

mounted devices [13] More recently MEMS temperature sensors have been developed

that consist of micro-machined semiconductor material that undergoes structural

deformation with temperature changes A common implementation is a bent beam

structure whose temperature induced deflection δ is expressed as [14]

( ) 3

2 2

sin12 cos sinz

EA T LEI AEL

α βδ 2β β

∆=

+

(3)

Temperature sensors are well suited for localized fault monitoring and detection but can

only characterize a mechanical system consisting of multiple components through the use

of multiple sensors

9

213 ACOUSTIC EMISSION SENSORS

Acoustic emission (AE) sensors have been used to characterize wear in machine

tools and monitor bearing and gear problems in centrifugal pumps [1516] First

developed as a Non-Destructive Testing (NDT) technique to detect cracks in civil

structures these sensors detect acoustic emissions generated by the release of vibration

waves in a crystalline lattice due to plastic deformation or crack growth [16]

Measurements are made using piezoelectric transducers with high natural frequencies

100 kHz to 1 MHz to capture the ultrasonic AE emissions An AE sensor is useful as it

has the ability to detect subsurface cracks in gear teeth or bearings before appearing on

the surface causing further damage [17] More recently MEMS acoustic sensors have

been developed and one design includes multiple transducers on a single substrate

which each detect acoustic emission energy at different frequencies [18] This helps

distinguish spurious acoustic emissions arising from impact and friction from those

arising from plastic deformation When compared to typical piezoelectric sensors the

MEMS devices have lower sensitivities and fail to detect some acoustic emissions

[1819] In addition the acoustic emission signal suffers severe attenuation as is crosses

various interfaces such as a gearbox or bearing casings In one experiment consisting of

a pinion gear and an associated bearing a 44dB attenuation was seen between an AE

sensor placed directly on the pinion to one placed on the bearing casing and in some

cases intermediate loss of the signal was observed [20]

10

22 MEMS VIBRATION SENSORS

This thesis focuses on the development of a MEMS vibration sensor for fault

detection as the vibration signature of a mechanical system has the potential to give an

overview of the entire system Many types of MEMS vibration sensors exist with

different operating principals based on resonance the piezoelectric effect and

displacement variation

221 RESONANT SENSORS

Resonant sensors are based on the resonant frequency shift of a fixed-fixed micro-

beam that is excited into resonance using electrostatic thermal or piezoelectric methods

Under an applied axial load there is a shift in the resonant frequency which is expressed

as [21]

2 2

2

1 12 1

nn n

EI Fl2

fl A Eα γ

π ρ= +

I

(4)

where l is the beam length E is Youngrsquos Modulus I is the moment of inertia ρ is the

density A is the cross section area F is the applied axial force αn and γn are mode

dependant coefficients and are 4730 and 0295 for mode 1 in the case discussed The

axial force F is generated by a proof mass attached to one beam end which moves in

response to an external acceleration The performance of resonant sensors are dependant

on the quality factor Q which is the ratio of total energy stored in the system to the

energy lost per cycle and determines the sharpness and amplitude of the resonance peak

A high Q resonator develops a peak that is easily distinguishable in a phase locked loop

control circuit and allows for improved performance and resolution [3] A resonator Q

value is lowered by energy losses to the surrounding fluid attached supports and

11

material Losses to the fluid surrounding the MEMS structure is the dominant loss

mechanism for MEMS sensors operated in air and can be greatly reduced by vacuum

packaging which can be accomplished using glass-silicon anodic bonding [22] In

addition to increasing fabrication cost the long term stability of the vacuum degrades

with ultimate device failure stemming from leakage through micro cracks and defects

and out-gassing [22] Support losses arise from the restoring forces they generate and

can be reduce by balanced designs such as a double ended tuning fork (DETF) or

operation in higher modes [3] In addition the sensitivity of a resonant sensor can be

increased by the addition of micro-levers to increase the axial force In one such design a

two stage micro-lever increased the sensitivity by an order of magnitude over a more

conventional single lever arm [2324]

222 PIEZOELECTRIC SENSORS

Piezoelectric materials produce a charge in response to an applied force This is an

intrinsic effect in materials such as Zinc Oxide (ZnO) whereas materials such as Barium

Titanate or Lead Zirconate Titanate (PZT) need to be poled by placing the material in a

strong electric field at an elevated temperature [3] The piezoelectric effect is expressed

mathematical by [25]

i ij j ik kD d T Eε= + (5)

where D represents the electrical displacement (Cm2) d is the charge coefficient (CN)

T is the applied stress matrix (Nm2) ε is the permittivity matrix (Fm) and E is the

electric field matrix (Vm) The subscripts i=123 j=123456 and k=123 represent

the direction to which the physical property is related For micro-sensors poled in the out

of plane direction only the D3 term is relevant and if fabricated using ZnO or PZT only

12

the charge coefficients d31 d32 and d33 are relevant [25] The piezoelectric material is

commonly deposited as a thin film on the surface of compliant structures in bulk micro-

machined sensors or incorporated as a layer in surface micro-machining on a similar

structure [26] However the inclusion of multiple materials introduces large temperature

variations causing out of plane bending especially in surface micro-machined devices

[2627] Poled piezoelectric materials could be depolarized which serves to reduce the

piezoelectric affect in the material This can occur if the sensor is exposed to a strong

electric field of the opposite polarity high temperatures in excess of the Curie point or

high mechanical stress Also the piezoelectric effect reduces over time an effect that can

be reduced with the addition of composite elements [28] Foundry processes such as

PolyMUMPS and SoiMUMPS do not incorporate piezoelectric layers thus sensors using

this method of transduction must use a customized process The piezoelectric sensor

outlined in [26] uses polysilicon surface micro-machining similar to PolyMUMPS

however it incorporates a ZnO layer that is not available in the foundry process

223 DISPLACEMENT VARIATION SENSORS

Displacement variation sensors consist of a proof mass connected to a compliant

micro-flexure The movement of the proof mass is sensed using optical electron

tunneling or more commonly capacitive methods A recent optical accelerometer

utilized a novel multilayer nano-grating for in-plane displacement sensing of the proof

mass [29] For small space variations between the nano-gratings a large change in the

optical reflection amplitude of the grating was observed [30] Even though optical

sensors offer resolutions approaching the Brownian noise limit they are not available in

small packages that are easily placed in space constricted areas for a low cost [29]

13

The electron tunneling effect is observed when the proximity between two electrodes

one flexible and the other fixed is on the order of 10Ǻ [31] If the two electrodes are

biased with voltage VB a tunneling current It flows between the gap [31]

exp( )t B I tgI V xαinfin sdot minus Φ (6)

where α1 is a constant and equal to 1025 Ǻ-1 eV-frac12 Ф is the effective height of the

tunneling barrier and xtg is the tunneling gap When the flexible electrode moves in

response to acceleration a feedback controller maintains the original distance while

determining the applied acceleration Electron tunneling sensors are able to sense

accelerations in the microg-ng range due to the exponential relation between tunneling current

and gap However these highly sensitive devices are susceptible to thermo-mechanical

noise which could be reduced by operation at low pressure [32] It is also important for

the two electrodes to be coated with metal which is difficult to achieve with foundry

fabrication

Capacitive sensors are based on the parallel plate capacitor and are implemented

using inter-digitated comb fingers in an in-plane lateral or transverse drive or in an out of

plane parallel plate drive [33] These sensors have the advantage of high sensitivity low

power consumption small package low temperature dependence and are easily

integrated with existing foundry fabrication processes [34] The lateral drive offers a fully

linear response but low sensitivity whereas the transverse and parallel plate drives

achieve high sensitivity if gap distances are of a few microns and large capacitance areas

are utilized However both drives have a non-linear response and are prone to pull-in

Also if these sensors are fabricated using a surface micromachining process such as

PolyMUMPS where the ground plane is a few microns above the structural layer they

14

are prone to electrical failure due to shorting This was observed in an experiment where

similar sensors were placed in a vibrating environment with a peak of 120g and shorting

between the structural layer and ground plane was found to be the predominate mode of

failure [35] However with certain fabrication processes such as SoiMUMPS this issue

can be neglected as the substrate material below any moving component is removed

Most capacitance sensors are fabricated using an integrated CMOS (Complementary

Metal Oxide Semiconductor)-MEMS technique where the MEMS sensor is fabricated

before during or after the CMOS circuit fabrication [36] Although these sensors offer

high-sensitivity and small parasitic capacitance they suffer from in-plane and out of

plane curling of the beam which reduces the capacitance between adjacent comb fingers

[37] In addition multiple processing steps are required for device fabrication

More than the other transduction methods MEMS capacitive displacement sensors are

attractive for practical implementation in an automobile and therefore is the selected

method for the vibration sensor presented in this work Their low temperature dependence

is an ideal characteristic for an automobile environment that faces varying temperature If

implemented with off-chip readout circuitry they require a standard foundry fabrication

process which provides a strong platform for mass production and low unit cost

3 SENSOR STRUCTURE

This chapter presents the physical structure of the three MEMS vibration sensors

which includes the displacement sensing capacitive comb drive and the micro-

flexureproof mass system First the capacitive comb drive and the three variations

implemented the transverse drive lateral drive and differential tri-plate drive are

discussed Next the mechanical structure is presented and includes the micro-flexure

selection quality factor and equations of motion Finally instability that places a limit

on the operating range of the sensor as a result of a gap closing capacitive comb drive is

discussed

31 SENSING

Capacitive sensors are based on the parallel plate capacitor and are commonly

implemented using inter-digitated comb fingers A capacitance can be realized if a voltage

is placed between two closely spaced plates If the fringing electric field is neglected the

capacitance of the parallel plate capacitor Cpp (F) is expressed as [28]

air pppp

k AC

dε

= (7)

where kair=1 is the dielectric constant of air at atmospheric pressure ε=88542times10-12 F m-1

is the permittivity of free space App is the overlapping area (m2) and d is the distance

between the two plates (m)

15

16

Figure 3-1 An inter-digitated comb drive where changes in capacitance are generated by either changes in gap distance x0 or in the overlap area y0timest

Figure 3-1

For a set of inter-digitated comb fingers or comb drive shown in the total

capacitance Ctotal is expressed as [28]

0 0

0 0

( ) ( )( ) (Total cy y y yC N t

)x x x xε

+ += + + minus

(8)

where N is the number of comb fingers on the movable comb y0 is the initial comb finger

overlap length (m) tc is the out of plane comb finger thickness (m) and x0 is the inital

finger gap (m) In this work the comb drive is implemented in three arrangements

transverse lateral and differential tri-plate In a transverse comb drive changes in

capacitance are generated by transverse movements along the x-axis resulting in changes

to the gap x0 between the movable and fixed comb fingers The nonlinear relation

between total capacitance CTotalT and displacement is expressed as

17

00 0

1 1( ) (Total TC N ty

)x x x xε

= + minus +

(9)

Although the drive sensitivity part is not constant an average sensitivity can be

assumed for small displacements [38]

Total TC xpart

In a lateral comb drive changes in capacitance are generated by lateral movements

along the y-axis resulting in changes to the overlapping area y0timestc The total capacitance

CTotalL between all comb fingers is expressed as

00

2 ( )Total LN tC yx

yε= +

(10)

The sensitivity of the lateral comb drive SL is expressed as

0

2Total LL

C N tSy x

εpart= =

part

(11)

The lateral comb drive was first introduced by W C Tang T H Nguyen and R T

Howe [33] in an electrostatic resonator and has since been used numerous times in both

drive and sense configurations It has the advantage of a fully linear response and

negligible instability but offers low sensitivity The sensitivity for the lateral drive L1

implemented in this work is 0038 pF microm-1

A differential tri-plate drive shown in Figure 3-2 which is a variation of the

transverse drive is implemented to improve sensitivity and reduce noise [39]

18

Figure 3-2 An inter-digitated comb drive in a differential tri-plate arrangement where x01 x02 and x03 are the gap distances and x01= x02ltlt x03 to improve sensitivity

In the tri-plate drive the at rest gap distances between the two capacitors formed by any

movable finger is not equivalent For a differential arrangement insuring that x01=

x02ltltx03 the linearity for small displacements (less than 1 microm) is enhanced over the case

of a single ended arrangement where x01=x02=x03=x0 which is equivalent to the

transverse drive This is shown graphically in Figure 3-3 where x01= x02=x03=2 microm in

the lower curve and x01= x02= 2 microm and x03= 15 microm in the upper curve A substantial

improvement in linearity is attained with the tri-plate drive

19

0

005

01

015

02

025

03

035

0 02 04 06 08 1

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-3 Comparison of the effect of gap distances on linearity for two case lsquorsquo represents case 1 where the fingers in are implemented in a differential tri-plate arrangement and x01= x02= 2 microm and x03= 15 microm lsquorsquo represents the case 2 where the fingers are implemented in a single ended arrangement and x01= x02=x03=2 microm

The tri-plate drive is also implemented in a differential arrangement where the output

is the change in capacitance rather than the total capacitance resulting from an external

excitation This reduces errors cause by off-axis excitation and common mode noise The

output of the differential tri-plate drive is expressed as [39]

Total Top Total BottomC C C∆ = minus where

003 01

1 1( ) (Total TopC N ty

)x x x xε

= +

minus +

(12)

20

002 03

1 1( ) (Total BottomC N ty

)x x x xε

= +

minus +

A graphic comparison of the capacitancedisplacement relation for sensors T1 and T2 is

shown in Figure 3-4 For these gap closing drives it is assumed that for displacements

less that 1 microm the relation is linear indicated by the dashed line above each curve The

sensitivity within this region is approximately 0046 pF microm-1 and 0253 pF microm-1 for T1

and T2 respectively The order of magnitude improvement is attributed to the differential

tri-plate arrangement and thicker comb finger

0

02

04

06

08

1

12

14

16

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-4 lsquorsquo represents the output of the differential tri-plate drive while lsquorsquo represents the output of the transverse drive The dashed line represents the linear approximation for displacements less that 1 microm For the purpose of comparison the output of the transverse drive was equalized to zero at zero displacement

21

In this work the transverse drive is first implemented in sensor T1 using the

PolyMUMPS fabrication process Then in an effort to linearize the sensor response the

lateral drive is used in L1 using SoiMUMPS The reduction in drive sensitivity due to the

lateral drive is further reduced by increasing the gap distance to 3 microm which is

recommended by the foundry to ensure a successful device These reductions are

balanced by increasing the number of comb fingers N and selecting a substrate thickness

that is 125 times large than T1 The third device T2 is fabricated using SoiMUMPS and

uses the tri-plate differential drive This design capitalizes on the increase in sensitivity of

a gap closing drive and larger substrate thickness attaining an order of magnitude

increase in sensitivity over T1 and L1 The specifications for each drive along with the

successive improvement in drive sensitivity are outlined in Table 3-1

Table 3-1 Specifications for each capacitive comb drive implemented in this work T1 L1 and T2

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Comb Finger Thickness tc (microm) 2 25 25 Gap Distance x0 (microm) 2 3 2 Number of Fingers N 200 256 30 Overlap length y0 (microm) 45 10 70 Drive Sensitivity (pF microm-1) 0046 0038 0273

32 MECHANICAL ANALYSIS

321 MICRO-FLEXURE SELECTION

The sensor structure consists of a proof mass connected to a compliant micro-flexure

that itself is anchored to the sensor substrate The proof mass displaces in response to an

22

applied acceleration and this displacement is sensed using the capacitive comb drive

attached to the proof mass

Proof mass

Folded beam micro-flexure

Anchor

Figure 3-5 Sensor Structure

This section focuses on micro-flexure selection as it should exhibit a large linear

displacement range and have a high lateral stiffness ratio to reduce cross-axis sensitivity

A number of micro-flexures have been used in MEMS research these include the

clamped-clamped crab-leg and folded beam shown in Figure 3-6-through 3-7

The clamped-clamped flexure exhibits a high stiffness ratio however for large lateral

displacements extensional axial forces buildup within the beams resulting in a non-linear

force-displacement relation [40] Within the linear regime the stiffness of the clamped-

clamped flexure is expressed as [40]

3

3b

yEt wk

L=

(13)

23

where E is the Youngrsquos Modulus of the beam material (N m-2) tb is the beam out of

plane thickness (m) w is the beam width (m) and L is the beam length (m) The stiffness

ratio is expressed as [40]

2x

y

k Lk w

=

(14)

If L is two orders of magnitude greater than w this results in a ratio of approximately

10000 The linear displacement range is determined using Finite Element Analysis (FEA)

of a clamped-clamped flexure with L=387 microm w=4 microm and tb=2 microm The result of this

simulation is plotted in and compared to a structure with constant stiffness ky

= 035 N m-1 predicted by Eq (13) The linear relation is valid for small deflections that

are less than 2 microm In addition this micro-flexure is susceptible to buckling as a result of

a compressive residual stress that may arise due to the fabrication process [41]

Figure 3-9

In the crab leg design the thigh section is meant to relieve the extensional axial

stress postponing the occurrence of non-linearity until greater lateral displacement [42]

The stiffness of the crab leg flexure is expressed as [43]

3

44

b s s ty

s s t

Et w L LkL L L

κκ

+= +

(15)

where κ=IsIt=(wswt)3 is a ratio of the second moment of inertia of the shin (s) and thigh

(t) ws and wt represent the width and Ls and Lt represent the length of the shin and thigh

regions respectively The stiffness ratio is expressed as [43]

34

4t s s tx

y s t s

w L L Lkk w L L L

αα t

+= +

(16)

The linear displacement range of a crab leg flexure with Ls=285 microm Lt=30 microm ws=wt=3

microm and tb=2 microm is determined using FEA The results are plotted in Figure 3-9 and

compared to a structure of constant stiffness ky=035 N m-1 predicted by Eq (15) The

24

non-linearity is greatly reduced when compared to the clamp-clamped flexure with Eq

accurately predicting the deflection until 4 microm However the flexure is more

susceptible to off-axis deflection as there is an order of magnitude reduction in the

stiffness ratio as compared to the clamped-clamped flexure

(15)

The folded beam flexure stands out as it is designed to exhibit a large linear

displacement range and to relieve residual stress due to the manufacturing process by

allowing the beams to expand and contract along the axial direction [40] It was first used

in an electrostatic resonator and has subsequently been used for electrostatic actuation in

a magnetic hard drives and micro-positioning [3344] If the truss section is rigid

ItrussgtgtIbeam then the stiffness of the flexure is expressed as [4043]

3

32b

yEt wk

L=

(17)

The stiffness ratio is expressed as [40]

2x

y

k Lk w

=

(18)

Figure 3-6 Clamped-clamped flexure

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

9

213 ACOUSTIC EMISSION SENSORS

Acoustic emission (AE) sensors have been used to characterize wear in machine

tools and monitor bearing and gear problems in centrifugal pumps [1516] First

developed as a Non-Destructive Testing (NDT) technique to detect cracks in civil

structures these sensors detect acoustic emissions generated by the release of vibration

waves in a crystalline lattice due to plastic deformation or crack growth [16]

Measurements are made using piezoelectric transducers with high natural frequencies

100 kHz to 1 MHz to capture the ultrasonic AE emissions An AE sensor is useful as it

has the ability to detect subsurface cracks in gear teeth or bearings before appearing on

the surface causing further damage [17] More recently MEMS acoustic sensors have

been developed and one design includes multiple transducers on a single substrate

which each detect acoustic emission energy at different frequencies [18] This helps

distinguish spurious acoustic emissions arising from impact and friction from those

arising from plastic deformation When compared to typical piezoelectric sensors the

MEMS devices have lower sensitivities and fail to detect some acoustic emissions

[1819] In addition the acoustic emission signal suffers severe attenuation as is crosses

various interfaces such as a gearbox or bearing casings In one experiment consisting of

a pinion gear and an associated bearing a 44dB attenuation was seen between an AE

sensor placed directly on the pinion to one placed on the bearing casing and in some

cases intermediate loss of the signal was observed [20]

10

22 MEMS VIBRATION SENSORS

This thesis focuses on the development of a MEMS vibration sensor for fault

detection as the vibration signature of a mechanical system has the potential to give an

overview of the entire system Many types of MEMS vibration sensors exist with

different operating principals based on resonance the piezoelectric effect and

displacement variation

221 RESONANT SENSORS

Resonant sensors are based on the resonant frequency shift of a fixed-fixed micro-

beam that is excited into resonance using electrostatic thermal or piezoelectric methods

Under an applied axial load there is a shift in the resonant frequency which is expressed

as [21]

2 2

2

1 12 1

nn n

EI Fl2

fl A Eα γ

π ρ= +

I

(4)

where l is the beam length E is Youngrsquos Modulus I is the moment of inertia ρ is the

density A is the cross section area F is the applied axial force αn and γn are mode

dependant coefficients and are 4730 and 0295 for mode 1 in the case discussed The

axial force F is generated by a proof mass attached to one beam end which moves in

response to an external acceleration The performance of resonant sensors are dependant

on the quality factor Q which is the ratio of total energy stored in the system to the

energy lost per cycle and determines the sharpness and amplitude of the resonance peak

A high Q resonator develops a peak that is easily distinguishable in a phase locked loop

control circuit and allows for improved performance and resolution [3] A resonator Q

value is lowered by energy losses to the surrounding fluid attached supports and

11

material Losses to the fluid surrounding the MEMS structure is the dominant loss

mechanism for MEMS sensors operated in air and can be greatly reduced by vacuum

packaging which can be accomplished using glass-silicon anodic bonding [22] In

addition to increasing fabrication cost the long term stability of the vacuum degrades

with ultimate device failure stemming from leakage through micro cracks and defects

and out-gassing [22] Support losses arise from the restoring forces they generate and

can be reduce by balanced designs such as a double ended tuning fork (DETF) or

operation in higher modes [3] In addition the sensitivity of a resonant sensor can be

increased by the addition of micro-levers to increase the axial force In one such design a

two stage micro-lever increased the sensitivity by an order of magnitude over a more

conventional single lever arm [2324]

222 PIEZOELECTRIC SENSORS

Piezoelectric materials produce a charge in response to an applied force This is an

intrinsic effect in materials such as Zinc Oxide (ZnO) whereas materials such as Barium

Titanate or Lead Zirconate Titanate (PZT) need to be poled by placing the material in a

strong electric field at an elevated temperature [3] The piezoelectric effect is expressed

mathematical by [25]

i ij j ik kD d T Eε= + (5)

where D represents the electrical displacement (Cm2) d is the charge coefficient (CN)

T is the applied stress matrix (Nm2) ε is the permittivity matrix (Fm) and E is the

electric field matrix (Vm) The subscripts i=123 j=123456 and k=123 represent

the direction to which the physical property is related For micro-sensors poled in the out

of plane direction only the D3 term is relevant and if fabricated using ZnO or PZT only

12

the charge coefficients d31 d32 and d33 are relevant [25] The piezoelectric material is

commonly deposited as a thin film on the surface of compliant structures in bulk micro-

machined sensors or incorporated as a layer in surface micro-machining on a similar

structure [26] However the inclusion of multiple materials introduces large temperature

variations causing out of plane bending especially in surface micro-machined devices

[2627] Poled piezoelectric materials could be depolarized which serves to reduce the

piezoelectric affect in the material This can occur if the sensor is exposed to a strong

electric field of the opposite polarity high temperatures in excess of the Curie point or

high mechanical stress Also the piezoelectric effect reduces over time an effect that can

be reduced with the addition of composite elements [28] Foundry processes such as

PolyMUMPS and SoiMUMPS do not incorporate piezoelectric layers thus sensors using

this method of transduction must use a customized process The piezoelectric sensor

outlined in [26] uses polysilicon surface micro-machining similar to PolyMUMPS

however it incorporates a ZnO layer that is not available in the foundry process

223 DISPLACEMENT VARIATION SENSORS

Displacement variation sensors consist of a proof mass connected to a compliant

micro-flexure The movement of the proof mass is sensed using optical electron

tunneling or more commonly capacitive methods A recent optical accelerometer

utilized a novel multilayer nano-grating for in-plane displacement sensing of the proof

mass [29] For small space variations between the nano-gratings a large change in the

optical reflection amplitude of the grating was observed [30] Even though optical

sensors offer resolutions approaching the Brownian noise limit they are not available in

small packages that are easily placed in space constricted areas for a low cost [29]

13

The electron tunneling effect is observed when the proximity between two electrodes

one flexible and the other fixed is on the order of 10Ǻ [31] If the two electrodes are

biased with voltage VB a tunneling current It flows between the gap [31]

exp( )t B I tgI V xαinfin sdot minus Φ (6)

where α1 is a constant and equal to 1025 Ǻ-1 eV-frac12 Ф is the effective height of the

tunneling barrier and xtg is the tunneling gap When the flexible electrode moves in

response to acceleration a feedback controller maintains the original distance while

determining the applied acceleration Electron tunneling sensors are able to sense

accelerations in the microg-ng range due to the exponential relation between tunneling current

and gap However these highly sensitive devices are susceptible to thermo-mechanical

noise which could be reduced by operation at low pressure [32] It is also important for

the two electrodes to be coated with metal which is difficult to achieve with foundry

fabrication

Capacitive sensors are based on the parallel plate capacitor and are implemented

using inter-digitated comb fingers in an in-plane lateral or transverse drive or in an out of

plane parallel plate drive [33] These sensors have the advantage of high sensitivity low

power consumption small package low temperature dependence and are easily

integrated with existing foundry fabrication processes [34] The lateral drive offers a fully

linear response but low sensitivity whereas the transverse and parallel plate drives

achieve high sensitivity if gap distances are of a few microns and large capacitance areas

are utilized However both drives have a non-linear response and are prone to pull-in

Also if these sensors are fabricated using a surface micromachining process such as

PolyMUMPS where the ground plane is a few microns above the structural layer they

14

are prone to electrical failure due to shorting This was observed in an experiment where

similar sensors were placed in a vibrating environment with a peak of 120g and shorting

between the structural layer and ground plane was found to be the predominate mode of

failure [35] However with certain fabrication processes such as SoiMUMPS this issue

can be neglected as the substrate material below any moving component is removed

Most capacitance sensors are fabricated using an integrated CMOS (Complementary

Metal Oxide Semiconductor)-MEMS technique where the MEMS sensor is fabricated

before during or after the CMOS circuit fabrication [36] Although these sensors offer

high-sensitivity and small parasitic capacitance they suffer from in-plane and out of

plane curling of the beam which reduces the capacitance between adjacent comb fingers

[37] In addition multiple processing steps are required for device fabrication

More than the other transduction methods MEMS capacitive displacement sensors are

attractive for practical implementation in an automobile and therefore is the selected

method for the vibration sensor presented in this work Their low temperature dependence

is an ideal characteristic for an automobile environment that faces varying temperature If

implemented with off-chip readout circuitry they require a standard foundry fabrication

process which provides a strong platform for mass production and low unit cost

3 SENSOR STRUCTURE

This chapter presents the physical structure of the three MEMS vibration sensors

which includes the displacement sensing capacitive comb drive and the micro-

flexureproof mass system First the capacitive comb drive and the three variations

implemented the transverse drive lateral drive and differential tri-plate drive are

discussed Next the mechanical structure is presented and includes the micro-flexure

selection quality factor and equations of motion Finally instability that places a limit

on the operating range of the sensor as a result of a gap closing capacitive comb drive is

discussed

31 SENSING

Capacitive sensors are based on the parallel plate capacitor and are commonly

implemented using inter-digitated comb fingers A capacitance can be realized if a voltage

is placed between two closely spaced plates If the fringing electric field is neglected the

capacitance of the parallel plate capacitor Cpp (F) is expressed as [28]

air pppp

k AC

dε

= (7)

where kair=1 is the dielectric constant of air at atmospheric pressure ε=88542times10-12 F m-1

is the permittivity of free space App is the overlapping area (m2) and d is the distance

between the two plates (m)

15

16

Figure 3-1 An inter-digitated comb drive where changes in capacitance are generated by either changes in gap distance x0 or in the overlap area y0timest

Figure 3-1

For a set of inter-digitated comb fingers or comb drive shown in the total

capacitance Ctotal is expressed as [28]

0 0

0 0

( ) ( )( ) (Total cy y y yC N t

)x x x xε

+ += + + minus

(8)

where N is the number of comb fingers on the movable comb y0 is the initial comb finger

overlap length (m) tc is the out of plane comb finger thickness (m) and x0 is the inital

finger gap (m) In this work the comb drive is implemented in three arrangements

transverse lateral and differential tri-plate In a transverse comb drive changes in

capacitance are generated by transverse movements along the x-axis resulting in changes

to the gap x0 between the movable and fixed comb fingers The nonlinear relation

between total capacitance CTotalT and displacement is expressed as

17

00 0

1 1( ) (Total TC N ty

)x x x xε

= + minus +

(9)

Although the drive sensitivity part is not constant an average sensitivity can be

assumed for small displacements [38]

Total TC xpart

In a lateral comb drive changes in capacitance are generated by lateral movements

along the y-axis resulting in changes to the overlapping area y0timestc The total capacitance

CTotalL between all comb fingers is expressed as

00

2 ( )Total LN tC yx

yε= +

(10)

The sensitivity of the lateral comb drive SL is expressed as

0

2Total LL

C N tSy x

εpart= =

part

(11)

The lateral comb drive was first introduced by W C Tang T H Nguyen and R T

Howe [33] in an electrostatic resonator and has since been used numerous times in both

drive and sense configurations It has the advantage of a fully linear response and

negligible instability but offers low sensitivity The sensitivity for the lateral drive L1

implemented in this work is 0038 pF microm-1

A differential tri-plate drive shown in Figure 3-2 which is a variation of the

transverse drive is implemented to improve sensitivity and reduce noise [39]

18

Figure 3-2 An inter-digitated comb drive in a differential tri-plate arrangement where x01 x02 and x03 are the gap distances and x01= x02ltlt x03 to improve sensitivity

In the tri-plate drive the at rest gap distances between the two capacitors formed by any

movable finger is not equivalent For a differential arrangement insuring that x01=

x02ltltx03 the linearity for small displacements (less than 1 microm) is enhanced over the case

of a single ended arrangement where x01=x02=x03=x0 which is equivalent to the

transverse drive This is shown graphically in Figure 3-3 where x01= x02=x03=2 microm in

the lower curve and x01= x02= 2 microm and x03= 15 microm in the upper curve A substantial

improvement in linearity is attained with the tri-plate drive

19

0

005

01

015

02

025

03

035

0 02 04 06 08 1

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-3 Comparison of the effect of gap distances on linearity for two case lsquorsquo represents case 1 where the fingers in are implemented in a differential tri-plate arrangement and x01= x02= 2 microm and x03= 15 microm lsquorsquo represents the case 2 where the fingers are implemented in a single ended arrangement and x01= x02=x03=2 microm

The tri-plate drive is also implemented in a differential arrangement where the output

is the change in capacitance rather than the total capacitance resulting from an external

excitation This reduces errors cause by off-axis excitation and common mode noise The

output of the differential tri-plate drive is expressed as [39]

Total Top Total BottomC C C∆ = minus where

003 01

1 1( ) (Total TopC N ty

)x x x xε

= +

minus +

(12)

20

002 03

1 1( ) (Total BottomC N ty

)x x x xε

= +

minus +

A graphic comparison of the capacitancedisplacement relation for sensors T1 and T2 is

shown in Figure 3-4 For these gap closing drives it is assumed that for displacements

less that 1 microm the relation is linear indicated by the dashed line above each curve The

sensitivity within this region is approximately 0046 pF microm-1 and 0253 pF microm-1 for T1

and T2 respectively The order of magnitude improvement is attributed to the differential

tri-plate arrangement and thicker comb finger

0

02

04

06

08

1

12

14

16

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-4 lsquorsquo represents the output of the differential tri-plate drive while lsquorsquo represents the output of the transverse drive The dashed line represents the linear approximation for displacements less that 1 microm For the purpose of comparison the output of the transverse drive was equalized to zero at zero displacement

21

In this work the transverse drive is first implemented in sensor T1 using the

PolyMUMPS fabrication process Then in an effort to linearize the sensor response the

lateral drive is used in L1 using SoiMUMPS The reduction in drive sensitivity due to the

lateral drive is further reduced by increasing the gap distance to 3 microm which is

recommended by the foundry to ensure a successful device These reductions are

balanced by increasing the number of comb fingers N and selecting a substrate thickness

that is 125 times large than T1 The third device T2 is fabricated using SoiMUMPS and

uses the tri-plate differential drive This design capitalizes on the increase in sensitivity of

a gap closing drive and larger substrate thickness attaining an order of magnitude

increase in sensitivity over T1 and L1 The specifications for each drive along with the

successive improvement in drive sensitivity are outlined in Table 3-1

Table 3-1 Specifications for each capacitive comb drive implemented in this work T1 L1 and T2

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Comb Finger Thickness tc (microm) 2 25 25 Gap Distance x0 (microm) 2 3 2 Number of Fingers N 200 256 30 Overlap length y0 (microm) 45 10 70 Drive Sensitivity (pF microm-1) 0046 0038 0273

32 MECHANICAL ANALYSIS

321 MICRO-FLEXURE SELECTION

The sensor structure consists of a proof mass connected to a compliant micro-flexure

that itself is anchored to the sensor substrate The proof mass displaces in response to an

22

applied acceleration and this displacement is sensed using the capacitive comb drive

attached to the proof mass

Proof mass

Folded beam micro-flexure

Anchor

Figure 3-5 Sensor Structure

This section focuses on micro-flexure selection as it should exhibit a large linear

displacement range and have a high lateral stiffness ratio to reduce cross-axis sensitivity

A number of micro-flexures have been used in MEMS research these include the

clamped-clamped crab-leg and folded beam shown in Figure 3-6-through 3-7

The clamped-clamped flexure exhibits a high stiffness ratio however for large lateral

displacements extensional axial forces buildup within the beams resulting in a non-linear

force-displacement relation [40] Within the linear regime the stiffness of the clamped-

clamped flexure is expressed as [40]

3

3b

yEt wk

L=

(13)

23

where E is the Youngrsquos Modulus of the beam material (N m-2) tb is the beam out of

plane thickness (m) w is the beam width (m) and L is the beam length (m) The stiffness

ratio is expressed as [40]

2x

y

k Lk w

=

(14)

If L is two orders of magnitude greater than w this results in a ratio of approximately

10000 The linear displacement range is determined using Finite Element Analysis (FEA)

of a clamped-clamped flexure with L=387 microm w=4 microm and tb=2 microm The result of this

simulation is plotted in and compared to a structure with constant stiffness ky

= 035 N m-1 predicted by Eq (13) The linear relation is valid for small deflections that

are less than 2 microm In addition this micro-flexure is susceptible to buckling as a result of

a compressive residual stress that may arise due to the fabrication process [41]

Figure 3-9

In the crab leg design the thigh section is meant to relieve the extensional axial

stress postponing the occurrence of non-linearity until greater lateral displacement [42]

The stiffness of the crab leg flexure is expressed as [43]

3

44

b s s ty

s s t

Et w L LkL L L

κκ

+= +

(15)

where κ=IsIt=(wswt)3 is a ratio of the second moment of inertia of the shin (s) and thigh

(t) ws and wt represent the width and Ls and Lt represent the length of the shin and thigh

regions respectively The stiffness ratio is expressed as [43]

34

4t s s tx

y s t s

w L L Lkk w L L L

αα t

+= +

(16)

The linear displacement range of a crab leg flexure with Ls=285 microm Lt=30 microm ws=wt=3

microm and tb=2 microm is determined using FEA The results are plotted in Figure 3-9 and

compared to a structure of constant stiffness ky=035 N m-1 predicted by Eq (15) The

24

non-linearity is greatly reduced when compared to the clamp-clamped flexure with Eq

accurately predicting the deflection until 4 microm However the flexure is more

susceptible to off-axis deflection as there is an order of magnitude reduction in the

stiffness ratio as compared to the clamped-clamped flexure

(15)

The folded beam flexure stands out as it is designed to exhibit a large linear

displacement range and to relieve residual stress due to the manufacturing process by

allowing the beams to expand and contract along the axial direction [40] It was first used

in an electrostatic resonator and has subsequently been used for electrostatic actuation in

a magnetic hard drives and micro-positioning [3344] If the truss section is rigid

ItrussgtgtIbeam then the stiffness of the flexure is expressed as [4043]

3

32b

yEt wk

L=

(17)

The stiffness ratio is expressed as [40]

2x

y

k Lk w

=

(18)

Figure 3-6 Clamped-clamped flexure

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

10

22 MEMS VIBRATION SENSORS

This thesis focuses on the development of a MEMS vibration sensor for fault

detection as the vibration signature of a mechanical system has the potential to give an

overview of the entire system Many types of MEMS vibration sensors exist with

different operating principals based on resonance the piezoelectric effect and

displacement variation

221 RESONANT SENSORS

Resonant sensors are based on the resonant frequency shift of a fixed-fixed micro-

beam that is excited into resonance using electrostatic thermal or piezoelectric methods

Under an applied axial load there is a shift in the resonant frequency which is expressed

as [21]

2 2

2

1 12 1

nn n

EI Fl2

fl A Eα γ

π ρ= +

I

(4)

where l is the beam length E is Youngrsquos Modulus I is the moment of inertia ρ is the

density A is the cross section area F is the applied axial force αn and γn are mode

dependant coefficients and are 4730 and 0295 for mode 1 in the case discussed The

axial force F is generated by a proof mass attached to one beam end which moves in

response to an external acceleration The performance of resonant sensors are dependant

on the quality factor Q which is the ratio of total energy stored in the system to the

energy lost per cycle and determines the sharpness and amplitude of the resonance peak

A high Q resonator develops a peak that is easily distinguishable in a phase locked loop

control circuit and allows for improved performance and resolution [3] A resonator Q

value is lowered by energy losses to the surrounding fluid attached supports and

11

material Losses to the fluid surrounding the MEMS structure is the dominant loss

mechanism for MEMS sensors operated in air and can be greatly reduced by vacuum

packaging which can be accomplished using glass-silicon anodic bonding [22] In

addition to increasing fabrication cost the long term stability of the vacuum degrades

with ultimate device failure stemming from leakage through micro cracks and defects

and out-gassing [22] Support losses arise from the restoring forces they generate and

can be reduce by balanced designs such as a double ended tuning fork (DETF) or

operation in higher modes [3] In addition the sensitivity of a resonant sensor can be

increased by the addition of micro-levers to increase the axial force In one such design a

two stage micro-lever increased the sensitivity by an order of magnitude over a more

conventional single lever arm [2324]

222 PIEZOELECTRIC SENSORS

Piezoelectric materials produce a charge in response to an applied force This is an

intrinsic effect in materials such as Zinc Oxide (ZnO) whereas materials such as Barium

Titanate or Lead Zirconate Titanate (PZT) need to be poled by placing the material in a

strong electric field at an elevated temperature [3] The piezoelectric effect is expressed

mathematical by [25]

i ij j ik kD d T Eε= + (5)

where D represents the electrical displacement (Cm2) d is the charge coefficient (CN)

T is the applied stress matrix (Nm2) ε is the permittivity matrix (Fm) and E is the

electric field matrix (Vm) The subscripts i=123 j=123456 and k=123 represent

the direction to which the physical property is related For micro-sensors poled in the out

of plane direction only the D3 term is relevant and if fabricated using ZnO or PZT only

12

the charge coefficients d31 d32 and d33 are relevant [25] The piezoelectric material is

commonly deposited as a thin film on the surface of compliant structures in bulk micro-

machined sensors or incorporated as a layer in surface micro-machining on a similar

structure [26] However the inclusion of multiple materials introduces large temperature

variations causing out of plane bending especially in surface micro-machined devices

[2627] Poled piezoelectric materials could be depolarized which serves to reduce the

piezoelectric affect in the material This can occur if the sensor is exposed to a strong

electric field of the opposite polarity high temperatures in excess of the Curie point or

high mechanical stress Also the piezoelectric effect reduces over time an effect that can

be reduced with the addition of composite elements [28] Foundry processes such as

PolyMUMPS and SoiMUMPS do not incorporate piezoelectric layers thus sensors using

this method of transduction must use a customized process The piezoelectric sensor

outlined in [26] uses polysilicon surface micro-machining similar to PolyMUMPS

however it incorporates a ZnO layer that is not available in the foundry process

223 DISPLACEMENT VARIATION SENSORS

Displacement variation sensors consist of a proof mass connected to a compliant

micro-flexure The movement of the proof mass is sensed using optical electron

tunneling or more commonly capacitive methods A recent optical accelerometer

utilized a novel multilayer nano-grating for in-plane displacement sensing of the proof

mass [29] For small space variations between the nano-gratings a large change in the

optical reflection amplitude of the grating was observed [30] Even though optical

sensors offer resolutions approaching the Brownian noise limit they are not available in

small packages that are easily placed in space constricted areas for a low cost [29]

13

The electron tunneling effect is observed when the proximity between two electrodes

one flexible and the other fixed is on the order of 10Ǻ [31] If the two electrodes are

biased with voltage VB a tunneling current It flows between the gap [31]

exp( )t B I tgI V xαinfin sdot minus Φ (6)

where α1 is a constant and equal to 1025 Ǻ-1 eV-frac12 Ф is the effective height of the

tunneling barrier and xtg is the tunneling gap When the flexible electrode moves in

response to acceleration a feedback controller maintains the original distance while

determining the applied acceleration Electron tunneling sensors are able to sense

accelerations in the microg-ng range due to the exponential relation between tunneling current

and gap However these highly sensitive devices are susceptible to thermo-mechanical

noise which could be reduced by operation at low pressure [32] It is also important for

the two electrodes to be coated with metal which is difficult to achieve with foundry

fabrication

Capacitive sensors are based on the parallel plate capacitor and are implemented

using inter-digitated comb fingers in an in-plane lateral or transverse drive or in an out of

plane parallel plate drive [33] These sensors have the advantage of high sensitivity low

power consumption small package low temperature dependence and are easily

integrated with existing foundry fabrication processes [34] The lateral drive offers a fully

linear response but low sensitivity whereas the transverse and parallel plate drives

achieve high sensitivity if gap distances are of a few microns and large capacitance areas

are utilized However both drives have a non-linear response and are prone to pull-in

Also if these sensors are fabricated using a surface micromachining process such as

PolyMUMPS where the ground plane is a few microns above the structural layer they

14

are prone to electrical failure due to shorting This was observed in an experiment where

similar sensors were placed in a vibrating environment with a peak of 120g and shorting

between the structural layer and ground plane was found to be the predominate mode of

failure [35] However with certain fabrication processes such as SoiMUMPS this issue

can be neglected as the substrate material below any moving component is removed

Most capacitance sensors are fabricated using an integrated CMOS (Complementary

Metal Oxide Semiconductor)-MEMS technique where the MEMS sensor is fabricated

before during or after the CMOS circuit fabrication [36] Although these sensors offer

high-sensitivity and small parasitic capacitance they suffer from in-plane and out of

plane curling of the beam which reduces the capacitance between adjacent comb fingers

[37] In addition multiple processing steps are required for device fabrication

More than the other transduction methods MEMS capacitive displacement sensors are

attractive for practical implementation in an automobile and therefore is the selected

method for the vibration sensor presented in this work Their low temperature dependence

is an ideal characteristic for an automobile environment that faces varying temperature If

implemented with off-chip readout circuitry they require a standard foundry fabrication

process which provides a strong platform for mass production and low unit cost

3 SENSOR STRUCTURE

This chapter presents the physical structure of the three MEMS vibration sensors

which includes the displacement sensing capacitive comb drive and the micro-

flexureproof mass system First the capacitive comb drive and the three variations

implemented the transverse drive lateral drive and differential tri-plate drive are

discussed Next the mechanical structure is presented and includes the micro-flexure

selection quality factor and equations of motion Finally instability that places a limit

on the operating range of the sensor as a result of a gap closing capacitive comb drive is

discussed

31 SENSING

Capacitive sensors are based on the parallel plate capacitor and are commonly

implemented using inter-digitated comb fingers A capacitance can be realized if a voltage

is placed between two closely spaced plates If the fringing electric field is neglected the

capacitance of the parallel plate capacitor Cpp (F) is expressed as [28]

air pppp

k AC

dε

= (7)

where kair=1 is the dielectric constant of air at atmospheric pressure ε=88542times10-12 F m-1

is the permittivity of free space App is the overlapping area (m2) and d is the distance

between the two plates (m)

15

16

Figure 3-1 An inter-digitated comb drive where changes in capacitance are generated by either changes in gap distance x0 or in the overlap area y0timest

Figure 3-1

For a set of inter-digitated comb fingers or comb drive shown in the total

capacitance Ctotal is expressed as [28]

0 0

0 0

( ) ( )( ) (Total cy y y yC N t

)x x x xε

+ += + + minus

(8)

where N is the number of comb fingers on the movable comb y0 is the initial comb finger

overlap length (m) tc is the out of plane comb finger thickness (m) and x0 is the inital

finger gap (m) In this work the comb drive is implemented in three arrangements

transverse lateral and differential tri-plate In a transverse comb drive changes in

capacitance are generated by transverse movements along the x-axis resulting in changes

to the gap x0 between the movable and fixed comb fingers The nonlinear relation

between total capacitance CTotalT and displacement is expressed as

17

00 0

1 1( ) (Total TC N ty

)x x x xε

= + minus +

(9)

Although the drive sensitivity part is not constant an average sensitivity can be

assumed for small displacements [38]

Total TC xpart

In a lateral comb drive changes in capacitance are generated by lateral movements

along the y-axis resulting in changes to the overlapping area y0timestc The total capacitance

CTotalL between all comb fingers is expressed as

00

2 ( )Total LN tC yx

yε= +

(10)

The sensitivity of the lateral comb drive SL is expressed as

0

2Total LL

C N tSy x

εpart= =

part

(11)

The lateral comb drive was first introduced by W C Tang T H Nguyen and R T

Howe [33] in an electrostatic resonator and has since been used numerous times in both

drive and sense configurations It has the advantage of a fully linear response and

negligible instability but offers low sensitivity The sensitivity for the lateral drive L1

implemented in this work is 0038 pF microm-1

A differential tri-plate drive shown in Figure 3-2 which is a variation of the

transverse drive is implemented to improve sensitivity and reduce noise [39]

18

Figure 3-2 An inter-digitated comb drive in a differential tri-plate arrangement where x01 x02 and x03 are the gap distances and x01= x02ltlt x03 to improve sensitivity

In the tri-plate drive the at rest gap distances between the two capacitors formed by any

movable finger is not equivalent For a differential arrangement insuring that x01=

x02ltltx03 the linearity for small displacements (less than 1 microm) is enhanced over the case

of a single ended arrangement where x01=x02=x03=x0 which is equivalent to the

transverse drive This is shown graphically in Figure 3-3 where x01= x02=x03=2 microm in

the lower curve and x01= x02= 2 microm and x03= 15 microm in the upper curve A substantial

improvement in linearity is attained with the tri-plate drive

19

0

005

01

015

02

025

03

035

0 02 04 06 08 1

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-3 Comparison of the effect of gap distances on linearity for two case lsquorsquo represents case 1 where the fingers in are implemented in a differential tri-plate arrangement and x01= x02= 2 microm and x03= 15 microm lsquorsquo represents the case 2 where the fingers are implemented in a single ended arrangement and x01= x02=x03=2 microm

The tri-plate drive is also implemented in a differential arrangement where the output

is the change in capacitance rather than the total capacitance resulting from an external

excitation This reduces errors cause by off-axis excitation and common mode noise The

output of the differential tri-plate drive is expressed as [39]

Total Top Total BottomC C C∆ = minus where

003 01

1 1( ) (Total TopC N ty

)x x x xε

= +

minus +

(12)

20

002 03

1 1( ) (Total BottomC N ty

)x x x xε

= +

minus +

A graphic comparison of the capacitancedisplacement relation for sensors T1 and T2 is

shown in Figure 3-4 For these gap closing drives it is assumed that for displacements

less that 1 microm the relation is linear indicated by the dashed line above each curve The

sensitivity within this region is approximately 0046 pF microm-1 and 0253 pF microm-1 for T1

and T2 respectively The order of magnitude improvement is attributed to the differential

tri-plate arrangement and thicker comb finger

0

02

04

06

08

1

12

14

16

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-4 lsquorsquo represents the output of the differential tri-plate drive while lsquorsquo represents the output of the transverse drive The dashed line represents the linear approximation for displacements less that 1 microm For the purpose of comparison the output of the transverse drive was equalized to zero at zero displacement

21

In this work the transverse drive is first implemented in sensor T1 using the

PolyMUMPS fabrication process Then in an effort to linearize the sensor response the

lateral drive is used in L1 using SoiMUMPS The reduction in drive sensitivity due to the

lateral drive is further reduced by increasing the gap distance to 3 microm which is

recommended by the foundry to ensure a successful device These reductions are

balanced by increasing the number of comb fingers N and selecting a substrate thickness

that is 125 times large than T1 The third device T2 is fabricated using SoiMUMPS and

uses the tri-plate differential drive This design capitalizes on the increase in sensitivity of

a gap closing drive and larger substrate thickness attaining an order of magnitude

increase in sensitivity over T1 and L1 The specifications for each drive along with the

successive improvement in drive sensitivity are outlined in Table 3-1

Table 3-1 Specifications for each capacitive comb drive implemented in this work T1 L1 and T2

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Comb Finger Thickness tc (microm) 2 25 25 Gap Distance x0 (microm) 2 3 2 Number of Fingers N 200 256 30 Overlap length y0 (microm) 45 10 70 Drive Sensitivity (pF microm-1) 0046 0038 0273

32 MECHANICAL ANALYSIS

321 MICRO-FLEXURE SELECTION

The sensor structure consists of a proof mass connected to a compliant micro-flexure

that itself is anchored to the sensor substrate The proof mass displaces in response to an

22

applied acceleration and this displacement is sensed using the capacitive comb drive

attached to the proof mass

Proof mass

Folded beam micro-flexure

Anchor

Figure 3-5 Sensor Structure

This section focuses on micro-flexure selection as it should exhibit a large linear

displacement range and have a high lateral stiffness ratio to reduce cross-axis sensitivity

A number of micro-flexures have been used in MEMS research these include the

clamped-clamped crab-leg and folded beam shown in Figure 3-6-through 3-7

The clamped-clamped flexure exhibits a high stiffness ratio however for large lateral

displacements extensional axial forces buildup within the beams resulting in a non-linear

force-displacement relation [40] Within the linear regime the stiffness of the clamped-

clamped flexure is expressed as [40]

3

3b

yEt wk

L=

(13)

23

where E is the Youngrsquos Modulus of the beam material (N m-2) tb is the beam out of

plane thickness (m) w is the beam width (m) and L is the beam length (m) The stiffness

ratio is expressed as [40]

2x

y

k Lk w

=

(14)

If L is two orders of magnitude greater than w this results in a ratio of approximately

10000 The linear displacement range is determined using Finite Element Analysis (FEA)

of a clamped-clamped flexure with L=387 microm w=4 microm and tb=2 microm The result of this

simulation is plotted in and compared to a structure with constant stiffness ky

= 035 N m-1 predicted by Eq (13) The linear relation is valid for small deflections that

are less than 2 microm In addition this micro-flexure is susceptible to buckling as a result of

a compressive residual stress that may arise due to the fabrication process [41]

Figure 3-9

In the crab leg design the thigh section is meant to relieve the extensional axial

stress postponing the occurrence of non-linearity until greater lateral displacement [42]

The stiffness of the crab leg flexure is expressed as [43]

3

44

b s s ty

s s t

Et w L LkL L L

κκ

+= +

(15)

where κ=IsIt=(wswt)3 is a ratio of the second moment of inertia of the shin (s) and thigh

(t) ws and wt represent the width and Ls and Lt represent the length of the shin and thigh

regions respectively The stiffness ratio is expressed as [43]

34

4t s s tx

y s t s

w L L Lkk w L L L

αα t

+= +

(16)

The linear displacement range of a crab leg flexure with Ls=285 microm Lt=30 microm ws=wt=3

microm and tb=2 microm is determined using FEA The results are plotted in Figure 3-9 and

compared to a structure of constant stiffness ky=035 N m-1 predicted by Eq (15) The

24

non-linearity is greatly reduced when compared to the clamp-clamped flexure with Eq

accurately predicting the deflection until 4 microm However the flexure is more

susceptible to off-axis deflection as there is an order of magnitude reduction in the

stiffness ratio as compared to the clamped-clamped flexure

(15)

The folded beam flexure stands out as it is designed to exhibit a large linear

displacement range and to relieve residual stress due to the manufacturing process by

allowing the beams to expand and contract along the axial direction [40] It was first used

in an electrostatic resonator and has subsequently been used for electrostatic actuation in

a magnetic hard drives and micro-positioning [3344] If the truss section is rigid

ItrussgtgtIbeam then the stiffness of the flexure is expressed as [4043]

3

32b

yEt wk

L=

(17)

The stiffness ratio is expressed as [40]

2x

y

k Lk w

=

(18)

Figure 3-6 Clamped-clamped flexure

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

11

material Losses to the fluid surrounding the MEMS structure is the dominant loss

mechanism for MEMS sensors operated in air and can be greatly reduced by vacuum

packaging which can be accomplished using glass-silicon anodic bonding [22] In

addition to increasing fabrication cost the long term stability of the vacuum degrades

with ultimate device failure stemming from leakage through micro cracks and defects

and out-gassing [22] Support losses arise from the restoring forces they generate and

can be reduce by balanced designs such as a double ended tuning fork (DETF) or

operation in higher modes [3] In addition the sensitivity of a resonant sensor can be

increased by the addition of micro-levers to increase the axial force In one such design a

two stage micro-lever increased the sensitivity by an order of magnitude over a more

conventional single lever arm [2324]

222 PIEZOELECTRIC SENSORS

Piezoelectric materials produce a charge in response to an applied force This is an

intrinsic effect in materials such as Zinc Oxide (ZnO) whereas materials such as Barium

Titanate or Lead Zirconate Titanate (PZT) need to be poled by placing the material in a

strong electric field at an elevated temperature [3] The piezoelectric effect is expressed

mathematical by [25]

i ij j ik kD d T Eε= + (5)

where D represents the electrical displacement (Cm2) d is the charge coefficient (CN)

T is the applied stress matrix (Nm2) ε is the permittivity matrix (Fm) and E is the

electric field matrix (Vm) The subscripts i=123 j=123456 and k=123 represent

the direction to which the physical property is related For micro-sensors poled in the out

of plane direction only the D3 term is relevant and if fabricated using ZnO or PZT only

12

the charge coefficients d31 d32 and d33 are relevant [25] The piezoelectric material is

commonly deposited as a thin film on the surface of compliant structures in bulk micro-

machined sensors or incorporated as a layer in surface micro-machining on a similar

structure [26] However the inclusion of multiple materials introduces large temperature

variations causing out of plane bending especially in surface micro-machined devices

[2627] Poled piezoelectric materials could be depolarized which serves to reduce the

piezoelectric affect in the material This can occur if the sensor is exposed to a strong

electric field of the opposite polarity high temperatures in excess of the Curie point or

high mechanical stress Also the piezoelectric effect reduces over time an effect that can

be reduced with the addition of composite elements [28] Foundry processes such as

PolyMUMPS and SoiMUMPS do not incorporate piezoelectric layers thus sensors using

this method of transduction must use a customized process The piezoelectric sensor

outlined in [26] uses polysilicon surface micro-machining similar to PolyMUMPS

however it incorporates a ZnO layer that is not available in the foundry process

223 DISPLACEMENT VARIATION SENSORS

Displacement variation sensors consist of a proof mass connected to a compliant

micro-flexure The movement of the proof mass is sensed using optical electron

tunneling or more commonly capacitive methods A recent optical accelerometer

utilized a novel multilayer nano-grating for in-plane displacement sensing of the proof

mass [29] For small space variations between the nano-gratings a large change in the

optical reflection amplitude of the grating was observed [30] Even though optical

sensors offer resolutions approaching the Brownian noise limit they are not available in

small packages that are easily placed in space constricted areas for a low cost [29]

13

The electron tunneling effect is observed when the proximity between two electrodes

one flexible and the other fixed is on the order of 10Ǻ [31] If the two electrodes are

biased with voltage VB a tunneling current It flows between the gap [31]

exp( )t B I tgI V xαinfin sdot minus Φ (6)

where α1 is a constant and equal to 1025 Ǻ-1 eV-frac12 Ф is the effective height of the

tunneling barrier and xtg is the tunneling gap When the flexible electrode moves in

response to acceleration a feedback controller maintains the original distance while

determining the applied acceleration Electron tunneling sensors are able to sense

accelerations in the microg-ng range due to the exponential relation between tunneling current

and gap However these highly sensitive devices are susceptible to thermo-mechanical

noise which could be reduced by operation at low pressure [32] It is also important for

the two electrodes to be coated with metal which is difficult to achieve with foundry

fabrication

Capacitive sensors are based on the parallel plate capacitor and are implemented

using inter-digitated comb fingers in an in-plane lateral or transverse drive or in an out of

plane parallel plate drive [33] These sensors have the advantage of high sensitivity low

power consumption small package low temperature dependence and are easily

integrated with existing foundry fabrication processes [34] The lateral drive offers a fully

linear response but low sensitivity whereas the transverse and parallel plate drives

achieve high sensitivity if gap distances are of a few microns and large capacitance areas

are utilized However both drives have a non-linear response and are prone to pull-in

Also if these sensors are fabricated using a surface micromachining process such as

PolyMUMPS where the ground plane is a few microns above the structural layer they

14

are prone to electrical failure due to shorting This was observed in an experiment where

similar sensors were placed in a vibrating environment with a peak of 120g and shorting

between the structural layer and ground plane was found to be the predominate mode of

failure [35] However with certain fabrication processes such as SoiMUMPS this issue

can be neglected as the substrate material below any moving component is removed

Most capacitance sensors are fabricated using an integrated CMOS (Complementary

Metal Oxide Semiconductor)-MEMS technique where the MEMS sensor is fabricated

before during or after the CMOS circuit fabrication [36] Although these sensors offer

high-sensitivity and small parasitic capacitance they suffer from in-plane and out of

plane curling of the beam which reduces the capacitance between adjacent comb fingers

[37] In addition multiple processing steps are required for device fabrication

More than the other transduction methods MEMS capacitive displacement sensors are

attractive for practical implementation in an automobile and therefore is the selected

method for the vibration sensor presented in this work Their low temperature dependence

is an ideal characteristic for an automobile environment that faces varying temperature If

implemented with off-chip readout circuitry they require a standard foundry fabrication

process which provides a strong platform for mass production and low unit cost

3 SENSOR STRUCTURE

This chapter presents the physical structure of the three MEMS vibration sensors

which includes the displacement sensing capacitive comb drive and the micro-

flexureproof mass system First the capacitive comb drive and the three variations

implemented the transverse drive lateral drive and differential tri-plate drive are

discussed Next the mechanical structure is presented and includes the micro-flexure

selection quality factor and equations of motion Finally instability that places a limit

on the operating range of the sensor as a result of a gap closing capacitive comb drive is

discussed

31 SENSING

Capacitive sensors are based on the parallel plate capacitor and are commonly

implemented using inter-digitated comb fingers A capacitance can be realized if a voltage

is placed between two closely spaced plates If the fringing electric field is neglected the

capacitance of the parallel plate capacitor Cpp (F) is expressed as [28]

air pppp

k AC

dε

= (7)

where kair=1 is the dielectric constant of air at atmospheric pressure ε=88542times10-12 F m-1

is the permittivity of free space App is the overlapping area (m2) and d is the distance

between the two plates (m)

15

16

Figure 3-1 An inter-digitated comb drive where changes in capacitance are generated by either changes in gap distance x0 or in the overlap area y0timest

Figure 3-1

For a set of inter-digitated comb fingers or comb drive shown in the total

capacitance Ctotal is expressed as [28]

0 0

0 0

( ) ( )( ) (Total cy y y yC N t

)x x x xε

+ += + + minus

(8)

where N is the number of comb fingers on the movable comb y0 is the initial comb finger

overlap length (m) tc is the out of plane comb finger thickness (m) and x0 is the inital

finger gap (m) In this work the comb drive is implemented in three arrangements

transverse lateral and differential tri-plate In a transverse comb drive changes in

capacitance are generated by transverse movements along the x-axis resulting in changes

to the gap x0 between the movable and fixed comb fingers The nonlinear relation

between total capacitance CTotalT and displacement is expressed as

17

00 0

1 1( ) (Total TC N ty

)x x x xε

= + minus +

(9)

Although the drive sensitivity part is not constant an average sensitivity can be

assumed for small displacements [38]

Total TC xpart

In a lateral comb drive changes in capacitance are generated by lateral movements

along the y-axis resulting in changes to the overlapping area y0timestc The total capacitance

CTotalL between all comb fingers is expressed as

00

2 ( )Total LN tC yx

yε= +

(10)

The sensitivity of the lateral comb drive SL is expressed as

0

2Total LL

C N tSy x

εpart= =

part

(11)

The lateral comb drive was first introduced by W C Tang T H Nguyen and R T

Howe [33] in an electrostatic resonator and has since been used numerous times in both

drive and sense configurations It has the advantage of a fully linear response and

negligible instability but offers low sensitivity The sensitivity for the lateral drive L1

implemented in this work is 0038 pF microm-1

A differential tri-plate drive shown in Figure 3-2 which is a variation of the

transverse drive is implemented to improve sensitivity and reduce noise [39]

18

Figure 3-2 An inter-digitated comb drive in a differential tri-plate arrangement where x01 x02 and x03 are the gap distances and x01= x02ltlt x03 to improve sensitivity

In the tri-plate drive the at rest gap distances between the two capacitors formed by any

movable finger is not equivalent For a differential arrangement insuring that x01=

x02ltltx03 the linearity for small displacements (less than 1 microm) is enhanced over the case

of a single ended arrangement where x01=x02=x03=x0 which is equivalent to the

transverse drive This is shown graphically in Figure 3-3 where x01= x02=x03=2 microm in

the lower curve and x01= x02= 2 microm and x03= 15 microm in the upper curve A substantial

improvement in linearity is attained with the tri-plate drive

19

0

005

01

015

02

025

03

035

0 02 04 06 08 1

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-3 Comparison of the effect of gap distances on linearity for two case lsquorsquo represents case 1 where the fingers in are implemented in a differential tri-plate arrangement and x01= x02= 2 microm and x03= 15 microm lsquorsquo represents the case 2 where the fingers are implemented in a single ended arrangement and x01= x02=x03=2 microm

The tri-plate drive is also implemented in a differential arrangement where the output

is the change in capacitance rather than the total capacitance resulting from an external

excitation This reduces errors cause by off-axis excitation and common mode noise The

output of the differential tri-plate drive is expressed as [39]

Total Top Total BottomC C C∆ = minus where

003 01

1 1( ) (Total TopC N ty

)x x x xε

= +

minus +

(12)

20

002 03

1 1( ) (Total BottomC N ty

)x x x xε

= +

minus +

A graphic comparison of the capacitancedisplacement relation for sensors T1 and T2 is

shown in Figure 3-4 For these gap closing drives it is assumed that for displacements

less that 1 microm the relation is linear indicated by the dashed line above each curve The

sensitivity within this region is approximately 0046 pF microm-1 and 0253 pF microm-1 for T1

and T2 respectively The order of magnitude improvement is attributed to the differential

tri-plate arrangement and thicker comb finger

0

02

04

06

08

1

12

14

16

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-4 lsquorsquo represents the output of the differential tri-plate drive while lsquorsquo represents the output of the transverse drive The dashed line represents the linear approximation for displacements less that 1 microm For the purpose of comparison the output of the transverse drive was equalized to zero at zero displacement

21

In this work the transverse drive is first implemented in sensor T1 using the

PolyMUMPS fabrication process Then in an effort to linearize the sensor response the

lateral drive is used in L1 using SoiMUMPS The reduction in drive sensitivity due to the

lateral drive is further reduced by increasing the gap distance to 3 microm which is

recommended by the foundry to ensure a successful device These reductions are

balanced by increasing the number of comb fingers N and selecting a substrate thickness

that is 125 times large than T1 The third device T2 is fabricated using SoiMUMPS and

uses the tri-plate differential drive This design capitalizes on the increase in sensitivity of

a gap closing drive and larger substrate thickness attaining an order of magnitude

increase in sensitivity over T1 and L1 The specifications for each drive along with the

successive improvement in drive sensitivity are outlined in Table 3-1

Table 3-1 Specifications for each capacitive comb drive implemented in this work T1 L1 and T2

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Comb Finger Thickness tc (microm) 2 25 25 Gap Distance x0 (microm) 2 3 2 Number of Fingers N 200 256 30 Overlap length y0 (microm) 45 10 70 Drive Sensitivity (pF microm-1) 0046 0038 0273

32 MECHANICAL ANALYSIS

321 MICRO-FLEXURE SELECTION

The sensor structure consists of a proof mass connected to a compliant micro-flexure

that itself is anchored to the sensor substrate The proof mass displaces in response to an

22

applied acceleration and this displacement is sensed using the capacitive comb drive

attached to the proof mass

Proof mass

Folded beam micro-flexure

Anchor

Figure 3-5 Sensor Structure

This section focuses on micro-flexure selection as it should exhibit a large linear

displacement range and have a high lateral stiffness ratio to reduce cross-axis sensitivity

A number of micro-flexures have been used in MEMS research these include the

clamped-clamped crab-leg and folded beam shown in Figure 3-6-through 3-7

The clamped-clamped flexure exhibits a high stiffness ratio however for large lateral

displacements extensional axial forces buildup within the beams resulting in a non-linear

force-displacement relation [40] Within the linear regime the stiffness of the clamped-

clamped flexure is expressed as [40]

3

3b

yEt wk

L=

(13)

23

where E is the Youngrsquos Modulus of the beam material (N m-2) tb is the beam out of

plane thickness (m) w is the beam width (m) and L is the beam length (m) The stiffness

ratio is expressed as [40]

2x

y

k Lk w

=

(14)

If L is two orders of magnitude greater than w this results in a ratio of approximately

10000 The linear displacement range is determined using Finite Element Analysis (FEA)

of a clamped-clamped flexure with L=387 microm w=4 microm and tb=2 microm The result of this

simulation is plotted in and compared to a structure with constant stiffness ky

= 035 N m-1 predicted by Eq (13) The linear relation is valid for small deflections that

are less than 2 microm In addition this micro-flexure is susceptible to buckling as a result of

a compressive residual stress that may arise due to the fabrication process [41]

Figure 3-9

In the crab leg design the thigh section is meant to relieve the extensional axial

stress postponing the occurrence of non-linearity until greater lateral displacement [42]

The stiffness of the crab leg flexure is expressed as [43]

3

44

b s s ty

s s t

Et w L LkL L L

κκ

+= +

(15)

where κ=IsIt=(wswt)3 is a ratio of the second moment of inertia of the shin (s) and thigh

(t) ws and wt represent the width and Ls and Lt represent the length of the shin and thigh

regions respectively The stiffness ratio is expressed as [43]

34

4t s s tx

y s t s

w L L Lkk w L L L

αα t

+= +

(16)

The linear displacement range of a crab leg flexure with Ls=285 microm Lt=30 microm ws=wt=3

microm and tb=2 microm is determined using FEA The results are plotted in Figure 3-9 and

compared to a structure of constant stiffness ky=035 N m-1 predicted by Eq (15) The

24

non-linearity is greatly reduced when compared to the clamp-clamped flexure with Eq

accurately predicting the deflection until 4 microm However the flexure is more

susceptible to off-axis deflection as there is an order of magnitude reduction in the

stiffness ratio as compared to the clamped-clamped flexure

(15)

The folded beam flexure stands out as it is designed to exhibit a large linear

displacement range and to relieve residual stress due to the manufacturing process by

allowing the beams to expand and contract along the axial direction [40] It was first used

in an electrostatic resonator and has subsequently been used for electrostatic actuation in

a magnetic hard drives and micro-positioning [3344] If the truss section is rigid

ItrussgtgtIbeam then the stiffness of the flexure is expressed as [4043]

3

32b

yEt wk

L=

(17)

The stiffness ratio is expressed as [40]

2x

y

k Lk w

=

(18)

Figure 3-6 Clamped-clamped flexure

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

12

the charge coefficients d31 d32 and d33 are relevant [25] The piezoelectric material is

commonly deposited as a thin film on the surface of compliant structures in bulk micro-

machined sensors or incorporated as a layer in surface micro-machining on a similar

structure [26] However the inclusion of multiple materials introduces large temperature

variations causing out of plane bending especially in surface micro-machined devices

[2627] Poled piezoelectric materials could be depolarized which serves to reduce the

piezoelectric affect in the material This can occur if the sensor is exposed to a strong

electric field of the opposite polarity high temperatures in excess of the Curie point or

high mechanical stress Also the piezoelectric effect reduces over time an effect that can

be reduced with the addition of composite elements [28] Foundry processes such as

PolyMUMPS and SoiMUMPS do not incorporate piezoelectric layers thus sensors using

this method of transduction must use a customized process The piezoelectric sensor

outlined in [26] uses polysilicon surface micro-machining similar to PolyMUMPS

however it incorporates a ZnO layer that is not available in the foundry process

223 DISPLACEMENT VARIATION SENSORS

Displacement variation sensors consist of a proof mass connected to a compliant

micro-flexure The movement of the proof mass is sensed using optical electron

tunneling or more commonly capacitive methods A recent optical accelerometer

utilized a novel multilayer nano-grating for in-plane displacement sensing of the proof

mass [29] For small space variations between the nano-gratings a large change in the

optical reflection amplitude of the grating was observed [30] Even though optical

sensors offer resolutions approaching the Brownian noise limit they are not available in

small packages that are easily placed in space constricted areas for a low cost [29]

13

The electron tunneling effect is observed when the proximity between two electrodes

one flexible and the other fixed is on the order of 10Ǻ [31] If the two electrodes are

biased with voltage VB a tunneling current It flows between the gap [31]

exp( )t B I tgI V xαinfin sdot minus Φ (6)

where α1 is a constant and equal to 1025 Ǻ-1 eV-frac12 Ф is the effective height of the

tunneling barrier and xtg is the tunneling gap When the flexible electrode moves in

response to acceleration a feedback controller maintains the original distance while

determining the applied acceleration Electron tunneling sensors are able to sense

accelerations in the microg-ng range due to the exponential relation between tunneling current

and gap However these highly sensitive devices are susceptible to thermo-mechanical

noise which could be reduced by operation at low pressure [32] It is also important for

the two electrodes to be coated with metal which is difficult to achieve with foundry

fabrication

Capacitive sensors are based on the parallel plate capacitor and are implemented

using inter-digitated comb fingers in an in-plane lateral or transverse drive or in an out of

plane parallel plate drive [33] These sensors have the advantage of high sensitivity low

power consumption small package low temperature dependence and are easily

integrated with existing foundry fabrication processes [34] The lateral drive offers a fully

linear response but low sensitivity whereas the transverse and parallel plate drives

achieve high sensitivity if gap distances are of a few microns and large capacitance areas

are utilized However both drives have a non-linear response and are prone to pull-in

Also if these sensors are fabricated using a surface micromachining process such as

PolyMUMPS where the ground plane is a few microns above the structural layer they

14

are prone to electrical failure due to shorting This was observed in an experiment where

similar sensors were placed in a vibrating environment with a peak of 120g and shorting

between the structural layer and ground plane was found to be the predominate mode of

failure [35] However with certain fabrication processes such as SoiMUMPS this issue

can be neglected as the substrate material below any moving component is removed

Most capacitance sensors are fabricated using an integrated CMOS (Complementary

Metal Oxide Semiconductor)-MEMS technique where the MEMS sensor is fabricated

before during or after the CMOS circuit fabrication [36] Although these sensors offer

high-sensitivity and small parasitic capacitance they suffer from in-plane and out of

plane curling of the beam which reduces the capacitance between adjacent comb fingers

[37] In addition multiple processing steps are required for device fabrication

More than the other transduction methods MEMS capacitive displacement sensors are

attractive for practical implementation in an automobile and therefore is the selected

method for the vibration sensor presented in this work Their low temperature dependence

is an ideal characteristic for an automobile environment that faces varying temperature If

implemented with off-chip readout circuitry they require a standard foundry fabrication

process which provides a strong platform for mass production and low unit cost

3 SENSOR STRUCTURE

This chapter presents the physical structure of the three MEMS vibration sensors

which includes the displacement sensing capacitive comb drive and the micro-

flexureproof mass system First the capacitive comb drive and the three variations

implemented the transverse drive lateral drive and differential tri-plate drive are

discussed Next the mechanical structure is presented and includes the micro-flexure

selection quality factor and equations of motion Finally instability that places a limit

on the operating range of the sensor as a result of a gap closing capacitive comb drive is

discussed

31 SENSING

Capacitive sensors are based on the parallel plate capacitor and are commonly

implemented using inter-digitated comb fingers A capacitance can be realized if a voltage

is placed between two closely spaced plates If the fringing electric field is neglected the

capacitance of the parallel plate capacitor Cpp (F) is expressed as [28]

air pppp

k AC

dε

= (7)

where kair=1 is the dielectric constant of air at atmospheric pressure ε=88542times10-12 F m-1

is the permittivity of free space App is the overlapping area (m2) and d is the distance

between the two plates (m)

15

16

Figure 3-1 An inter-digitated comb drive where changes in capacitance are generated by either changes in gap distance x0 or in the overlap area y0timest

Figure 3-1

For a set of inter-digitated comb fingers or comb drive shown in the total

capacitance Ctotal is expressed as [28]

0 0

0 0

( ) ( )( ) (Total cy y y yC N t

)x x x xε

+ += + + minus

(8)

where N is the number of comb fingers on the movable comb y0 is the initial comb finger

overlap length (m) tc is the out of plane comb finger thickness (m) and x0 is the inital

finger gap (m) In this work the comb drive is implemented in three arrangements

transverse lateral and differential tri-plate In a transverse comb drive changes in

capacitance are generated by transverse movements along the x-axis resulting in changes

to the gap x0 between the movable and fixed comb fingers The nonlinear relation

between total capacitance CTotalT and displacement is expressed as

17

00 0

1 1( ) (Total TC N ty

)x x x xε

= + minus +

(9)

Although the drive sensitivity part is not constant an average sensitivity can be

assumed for small displacements [38]

Total TC xpart

In a lateral comb drive changes in capacitance are generated by lateral movements

along the y-axis resulting in changes to the overlapping area y0timestc The total capacitance

CTotalL between all comb fingers is expressed as

00

2 ( )Total LN tC yx

yε= +

(10)

The sensitivity of the lateral comb drive SL is expressed as

0

2Total LL

C N tSy x

εpart= =

part

(11)

The lateral comb drive was first introduced by W C Tang T H Nguyen and R T

Howe [33] in an electrostatic resonator and has since been used numerous times in both

drive and sense configurations It has the advantage of a fully linear response and

negligible instability but offers low sensitivity The sensitivity for the lateral drive L1

implemented in this work is 0038 pF microm-1

A differential tri-plate drive shown in Figure 3-2 which is a variation of the

transverse drive is implemented to improve sensitivity and reduce noise [39]

18

Figure 3-2 An inter-digitated comb drive in a differential tri-plate arrangement where x01 x02 and x03 are the gap distances and x01= x02ltlt x03 to improve sensitivity

In the tri-plate drive the at rest gap distances between the two capacitors formed by any

movable finger is not equivalent For a differential arrangement insuring that x01=

x02ltltx03 the linearity for small displacements (less than 1 microm) is enhanced over the case

of a single ended arrangement where x01=x02=x03=x0 which is equivalent to the

transverse drive This is shown graphically in Figure 3-3 where x01= x02=x03=2 microm in

the lower curve and x01= x02= 2 microm and x03= 15 microm in the upper curve A substantial

improvement in linearity is attained with the tri-plate drive

19

0

005

01

015

02

025

03

035

0 02 04 06 08 1

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-3 Comparison of the effect of gap distances on linearity for two case lsquorsquo represents case 1 where the fingers in are implemented in a differential tri-plate arrangement and x01= x02= 2 microm and x03= 15 microm lsquorsquo represents the case 2 where the fingers are implemented in a single ended arrangement and x01= x02=x03=2 microm

The tri-plate drive is also implemented in a differential arrangement where the output

is the change in capacitance rather than the total capacitance resulting from an external

excitation This reduces errors cause by off-axis excitation and common mode noise The

output of the differential tri-plate drive is expressed as [39]

Total Top Total BottomC C C∆ = minus where

003 01

1 1( ) (Total TopC N ty

)x x x xε

= +

minus +

(12)

20

002 03

1 1( ) (Total BottomC N ty

)x x x xε

= +

minus +

A graphic comparison of the capacitancedisplacement relation for sensors T1 and T2 is

shown in Figure 3-4 For these gap closing drives it is assumed that for displacements

less that 1 microm the relation is linear indicated by the dashed line above each curve The

sensitivity within this region is approximately 0046 pF microm-1 and 0253 pF microm-1 for T1

and T2 respectively The order of magnitude improvement is attributed to the differential

tri-plate arrangement and thicker comb finger

0

02

04

06

08

1

12

14

16

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-4 lsquorsquo represents the output of the differential tri-plate drive while lsquorsquo represents the output of the transverse drive The dashed line represents the linear approximation for displacements less that 1 microm For the purpose of comparison the output of the transverse drive was equalized to zero at zero displacement

21

In this work the transverse drive is first implemented in sensor T1 using the

PolyMUMPS fabrication process Then in an effort to linearize the sensor response the

lateral drive is used in L1 using SoiMUMPS The reduction in drive sensitivity due to the

lateral drive is further reduced by increasing the gap distance to 3 microm which is

recommended by the foundry to ensure a successful device These reductions are

balanced by increasing the number of comb fingers N and selecting a substrate thickness

that is 125 times large than T1 The third device T2 is fabricated using SoiMUMPS and

uses the tri-plate differential drive This design capitalizes on the increase in sensitivity of

a gap closing drive and larger substrate thickness attaining an order of magnitude

increase in sensitivity over T1 and L1 The specifications for each drive along with the

successive improvement in drive sensitivity are outlined in Table 3-1

Table 3-1 Specifications for each capacitive comb drive implemented in this work T1 L1 and T2

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Comb Finger Thickness tc (microm) 2 25 25 Gap Distance x0 (microm) 2 3 2 Number of Fingers N 200 256 30 Overlap length y0 (microm) 45 10 70 Drive Sensitivity (pF microm-1) 0046 0038 0273

32 MECHANICAL ANALYSIS

321 MICRO-FLEXURE SELECTION

The sensor structure consists of a proof mass connected to a compliant micro-flexure

that itself is anchored to the sensor substrate The proof mass displaces in response to an

22

applied acceleration and this displacement is sensed using the capacitive comb drive

attached to the proof mass

Proof mass

Folded beam micro-flexure

Anchor

Figure 3-5 Sensor Structure

This section focuses on micro-flexure selection as it should exhibit a large linear

displacement range and have a high lateral stiffness ratio to reduce cross-axis sensitivity

A number of micro-flexures have been used in MEMS research these include the

clamped-clamped crab-leg and folded beam shown in Figure 3-6-through 3-7

The clamped-clamped flexure exhibits a high stiffness ratio however for large lateral

displacements extensional axial forces buildup within the beams resulting in a non-linear

force-displacement relation [40] Within the linear regime the stiffness of the clamped-

clamped flexure is expressed as [40]

3

3b

yEt wk

L=

(13)

23

where E is the Youngrsquos Modulus of the beam material (N m-2) tb is the beam out of

plane thickness (m) w is the beam width (m) and L is the beam length (m) The stiffness

ratio is expressed as [40]

2x

y

k Lk w

=

(14)

If L is two orders of magnitude greater than w this results in a ratio of approximately

10000 The linear displacement range is determined using Finite Element Analysis (FEA)

of a clamped-clamped flexure with L=387 microm w=4 microm and tb=2 microm The result of this

simulation is plotted in and compared to a structure with constant stiffness ky

= 035 N m-1 predicted by Eq (13) The linear relation is valid for small deflections that

are less than 2 microm In addition this micro-flexure is susceptible to buckling as a result of

a compressive residual stress that may arise due to the fabrication process [41]

Figure 3-9

In the crab leg design the thigh section is meant to relieve the extensional axial

stress postponing the occurrence of non-linearity until greater lateral displacement [42]

The stiffness of the crab leg flexure is expressed as [43]

3

44

b s s ty

s s t

Et w L LkL L L

κκ

+= +

(15)

where κ=IsIt=(wswt)3 is a ratio of the second moment of inertia of the shin (s) and thigh

(t) ws and wt represent the width and Ls and Lt represent the length of the shin and thigh

regions respectively The stiffness ratio is expressed as [43]

34

4t s s tx

y s t s

w L L Lkk w L L L

αα t

+= +

(16)

The linear displacement range of a crab leg flexure with Ls=285 microm Lt=30 microm ws=wt=3

microm and tb=2 microm is determined using FEA The results are plotted in Figure 3-9 and

compared to a structure of constant stiffness ky=035 N m-1 predicted by Eq (15) The

24

non-linearity is greatly reduced when compared to the clamp-clamped flexure with Eq

accurately predicting the deflection until 4 microm However the flexure is more

susceptible to off-axis deflection as there is an order of magnitude reduction in the

stiffness ratio as compared to the clamped-clamped flexure

(15)

The folded beam flexure stands out as it is designed to exhibit a large linear

displacement range and to relieve residual stress due to the manufacturing process by

allowing the beams to expand and contract along the axial direction [40] It was first used

in an electrostatic resonator and has subsequently been used for electrostatic actuation in

a magnetic hard drives and micro-positioning [3344] If the truss section is rigid

ItrussgtgtIbeam then the stiffness of the flexure is expressed as [4043]

3

32b

yEt wk

L=

(17)

The stiffness ratio is expressed as [40]

2x

y

k Lk w

=

(18)

Figure 3-6 Clamped-clamped flexure

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

13

The electron tunneling effect is observed when the proximity between two electrodes

one flexible and the other fixed is on the order of 10Ǻ [31] If the two electrodes are

biased with voltage VB a tunneling current It flows between the gap [31]

exp( )t B I tgI V xαinfin sdot minus Φ (6)

where α1 is a constant and equal to 1025 Ǻ-1 eV-frac12 Ф is the effective height of the

tunneling barrier and xtg is the tunneling gap When the flexible electrode moves in

response to acceleration a feedback controller maintains the original distance while

determining the applied acceleration Electron tunneling sensors are able to sense

accelerations in the microg-ng range due to the exponential relation between tunneling current

and gap However these highly sensitive devices are susceptible to thermo-mechanical

noise which could be reduced by operation at low pressure [32] It is also important for

the two electrodes to be coated with metal which is difficult to achieve with foundry

fabrication

Capacitive sensors are based on the parallel plate capacitor and are implemented

using inter-digitated comb fingers in an in-plane lateral or transverse drive or in an out of

plane parallel plate drive [33] These sensors have the advantage of high sensitivity low

power consumption small package low temperature dependence and are easily

integrated with existing foundry fabrication processes [34] The lateral drive offers a fully

linear response but low sensitivity whereas the transverse and parallel plate drives

achieve high sensitivity if gap distances are of a few microns and large capacitance areas

are utilized However both drives have a non-linear response and are prone to pull-in

Also if these sensors are fabricated using a surface micromachining process such as

PolyMUMPS where the ground plane is a few microns above the structural layer they

14

are prone to electrical failure due to shorting This was observed in an experiment where

similar sensors were placed in a vibrating environment with a peak of 120g and shorting

between the structural layer and ground plane was found to be the predominate mode of

failure [35] However with certain fabrication processes such as SoiMUMPS this issue

can be neglected as the substrate material below any moving component is removed

Most capacitance sensors are fabricated using an integrated CMOS (Complementary

Metal Oxide Semiconductor)-MEMS technique where the MEMS sensor is fabricated

before during or after the CMOS circuit fabrication [36] Although these sensors offer

high-sensitivity and small parasitic capacitance they suffer from in-plane and out of

plane curling of the beam which reduces the capacitance between adjacent comb fingers

[37] In addition multiple processing steps are required for device fabrication

More than the other transduction methods MEMS capacitive displacement sensors are

attractive for practical implementation in an automobile and therefore is the selected

method for the vibration sensor presented in this work Their low temperature dependence

is an ideal characteristic for an automobile environment that faces varying temperature If

implemented with off-chip readout circuitry they require a standard foundry fabrication

process which provides a strong platform for mass production and low unit cost

3 SENSOR STRUCTURE

This chapter presents the physical structure of the three MEMS vibration sensors

which includes the displacement sensing capacitive comb drive and the micro-

flexureproof mass system First the capacitive comb drive and the three variations

implemented the transverse drive lateral drive and differential tri-plate drive are

discussed Next the mechanical structure is presented and includes the micro-flexure

selection quality factor and equations of motion Finally instability that places a limit

on the operating range of the sensor as a result of a gap closing capacitive comb drive is

discussed

31 SENSING

Capacitive sensors are based on the parallel plate capacitor and are commonly

implemented using inter-digitated comb fingers A capacitance can be realized if a voltage

is placed between two closely spaced plates If the fringing electric field is neglected the

capacitance of the parallel plate capacitor Cpp (F) is expressed as [28]

air pppp

k AC

dε

= (7)

where kair=1 is the dielectric constant of air at atmospheric pressure ε=88542times10-12 F m-1

is the permittivity of free space App is the overlapping area (m2) and d is the distance

between the two plates (m)

15

16

Figure 3-1 An inter-digitated comb drive where changes in capacitance are generated by either changes in gap distance x0 or in the overlap area y0timest

Figure 3-1

For a set of inter-digitated comb fingers or comb drive shown in the total

capacitance Ctotal is expressed as [28]

0 0

0 0

( ) ( )( ) (Total cy y y yC N t

)x x x xε

+ += + + minus

(8)

where N is the number of comb fingers on the movable comb y0 is the initial comb finger

overlap length (m) tc is the out of plane comb finger thickness (m) and x0 is the inital

finger gap (m) In this work the comb drive is implemented in three arrangements

transverse lateral and differential tri-plate In a transverse comb drive changes in

capacitance are generated by transverse movements along the x-axis resulting in changes

to the gap x0 between the movable and fixed comb fingers The nonlinear relation

between total capacitance CTotalT and displacement is expressed as

17

00 0

1 1( ) (Total TC N ty

)x x x xε

= + minus +

(9)

Although the drive sensitivity part is not constant an average sensitivity can be

assumed for small displacements [38]

Total TC xpart

In a lateral comb drive changes in capacitance are generated by lateral movements

along the y-axis resulting in changes to the overlapping area y0timestc The total capacitance

CTotalL between all comb fingers is expressed as

00

2 ( )Total LN tC yx

yε= +

(10)

The sensitivity of the lateral comb drive SL is expressed as

0

2Total LL

C N tSy x

εpart= =

part

(11)

The lateral comb drive was first introduced by W C Tang T H Nguyen and R T

Howe [33] in an electrostatic resonator and has since been used numerous times in both

drive and sense configurations It has the advantage of a fully linear response and

negligible instability but offers low sensitivity The sensitivity for the lateral drive L1

implemented in this work is 0038 pF microm-1

A differential tri-plate drive shown in Figure 3-2 which is a variation of the

transverse drive is implemented to improve sensitivity and reduce noise [39]

18

Figure 3-2 An inter-digitated comb drive in a differential tri-plate arrangement where x01 x02 and x03 are the gap distances and x01= x02ltlt x03 to improve sensitivity

In the tri-plate drive the at rest gap distances between the two capacitors formed by any

movable finger is not equivalent For a differential arrangement insuring that x01=

x02ltltx03 the linearity for small displacements (less than 1 microm) is enhanced over the case

of a single ended arrangement where x01=x02=x03=x0 which is equivalent to the

transverse drive This is shown graphically in Figure 3-3 where x01= x02=x03=2 microm in

the lower curve and x01= x02= 2 microm and x03= 15 microm in the upper curve A substantial

improvement in linearity is attained with the tri-plate drive

19

0

005

01

015

02

025

03

035

0 02 04 06 08 1

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-3 Comparison of the effect of gap distances on linearity for two case lsquorsquo represents case 1 where the fingers in are implemented in a differential tri-plate arrangement and x01= x02= 2 microm and x03= 15 microm lsquorsquo represents the case 2 where the fingers are implemented in a single ended arrangement and x01= x02=x03=2 microm

The tri-plate drive is also implemented in a differential arrangement where the output

is the change in capacitance rather than the total capacitance resulting from an external

excitation This reduces errors cause by off-axis excitation and common mode noise The

output of the differential tri-plate drive is expressed as [39]

Total Top Total BottomC C C∆ = minus where

003 01

1 1( ) (Total TopC N ty

)x x x xε

= +

minus +

(12)

20

002 03

1 1( ) (Total BottomC N ty

)x x x xε

= +

minus +

A graphic comparison of the capacitancedisplacement relation for sensors T1 and T2 is

shown in Figure 3-4 For these gap closing drives it is assumed that for displacements

less that 1 microm the relation is linear indicated by the dashed line above each curve The

sensitivity within this region is approximately 0046 pF microm-1 and 0253 pF microm-1 for T1

and T2 respectively The order of magnitude improvement is attributed to the differential

tri-plate arrangement and thicker comb finger

0

02

04

06

08

1

12

14

16

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-4 lsquorsquo represents the output of the differential tri-plate drive while lsquorsquo represents the output of the transverse drive The dashed line represents the linear approximation for displacements less that 1 microm For the purpose of comparison the output of the transverse drive was equalized to zero at zero displacement

21

In this work the transverse drive is first implemented in sensor T1 using the

PolyMUMPS fabrication process Then in an effort to linearize the sensor response the

lateral drive is used in L1 using SoiMUMPS The reduction in drive sensitivity due to the

lateral drive is further reduced by increasing the gap distance to 3 microm which is

recommended by the foundry to ensure a successful device These reductions are

balanced by increasing the number of comb fingers N and selecting a substrate thickness

that is 125 times large than T1 The third device T2 is fabricated using SoiMUMPS and

uses the tri-plate differential drive This design capitalizes on the increase in sensitivity of

a gap closing drive and larger substrate thickness attaining an order of magnitude

increase in sensitivity over T1 and L1 The specifications for each drive along with the

successive improvement in drive sensitivity are outlined in Table 3-1

Table 3-1 Specifications for each capacitive comb drive implemented in this work T1 L1 and T2

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Comb Finger Thickness tc (microm) 2 25 25 Gap Distance x0 (microm) 2 3 2 Number of Fingers N 200 256 30 Overlap length y0 (microm) 45 10 70 Drive Sensitivity (pF microm-1) 0046 0038 0273

32 MECHANICAL ANALYSIS

321 MICRO-FLEXURE SELECTION

The sensor structure consists of a proof mass connected to a compliant micro-flexure

that itself is anchored to the sensor substrate The proof mass displaces in response to an

22

applied acceleration and this displacement is sensed using the capacitive comb drive

attached to the proof mass

Proof mass

Folded beam micro-flexure

Anchor

Figure 3-5 Sensor Structure

This section focuses on micro-flexure selection as it should exhibit a large linear

displacement range and have a high lateral stiffness ratio to reduce cross-axis sensitivity

A number of micro-flexures have been used in MEMS research these include the

clamped-clamped crab-leg and folded beam shown in Figure 3-6-through 3-7

The clamped-clamped flexure exhibits a high stiffness ratio however for large lateral

displacements extensional axial forces buildup within the beams resulting in a non-linear

force-displacement relation [40] Within the linear regime the stiffness of the clamped-

clamped flexure is expressed as [40]

3

3b

yEt wk

L=

(13)

23

where E is the Youngrsquos Modulus of the beam material (N m-2) tb is the beam out of

plane thickness (m) w is the beam width (m) and L is the beam length (m) The stiffness

ratio is expressed as [40]

2x

y

k Lk w

=

(14)

If L is two orders of magnitude greater than w this results in a ratio of approximately

10000 The linear displacement range is determined using Finite Element Analysis (FEA)

of a clamped-clamped flexure with L=387 microm w=4 microm and tb=2 microm The result of this

simulation is plotted in and compared to a structure with constant stiffness ky

= 035 N m-1 predicted by Eq (13) The linear relation is valid for small deflections that

are less than 2 microm In addition this micro-flexure is susceptible to buckling as a result of

a compressive residual stress that may arise due to the fabrication process [41]

Figure 3-9

In the crab leg design the thigh section is meant to relieve the extensional axial

stress postponing the occurrence of non-linearity until greater lateral displacement [42]

The stiffness of the crab leg flexure is expressed as [43]

3

44

b s s ty

s s t

Et w L LkL L L

κκ

+= +

(15)

where κ=IsIt=(wswt)3 is a ratio of the second moment of inertia of the shin (s) and thigh

(t) ws and wt represent the width and Ls and Lt represent the length of the shin and thigh

regions respectively The stiffness ratio is expressed as [43]

34

4t s s tx

y s t s

w L L Lkk w L L L

αα t

+= +

(16)

The linear displacement range of a crab leg flexure with Ls=285 microm Lt=30 microm ws=wt=3

microm and tb=2 microm is determined using FEA The results are plotted in Figure 3-9 and

compared to a structure of constant stiffness ky=035 N m-1 predicted by Eq (15) The

24

non-linearity is greatly reduced when compared to the clamp-clamped flexure with Eq

accurately predicting the deflection until 4 microm However the flexure is more

susceptible to off-axis deflection as there is an order of magnitude reduction in the

stiffness ratio as compared to the clamped-clamped flexure

(15)

The folded beam flexure stands out as it is designed to exhibit a large linear

displacement range and to relieve residual stress due to the manufacturing process by

allowing the beams to expand and contract along the axial direction [40] It was first used

in an electrostatic resonator and has subsequently been used for electrostatic actuation in

a magnetic hard drives and micro-positioning [3344] If the truss section is rigid

ItrussgtgtIbeam then the stiffness of the flexure is expressed as [4043]

3

32b

yEt wk

L=

(17)

The stiffness ratio is expressed as [40]

2x

y

k Lk w

=

(18)

Figure 3-6 Clamped-clamped flexure

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

14

are prone to electrical failure due to shorting This was observed in an experiment where

similar sensors were placed in a vibrating environment with a peak of 120g and shorting

between the structural layer and ground plane was found to be the predominate mode of

failure [35] However with certain fabrication processes such as SoiMUMPS this issue

can be neglected as the substrate material below any moving component is removed

Most capacitance sensors are fabricated using an integrated CMOS (Complementary

Metal Oxide Semiconductor)-MEMS technique where the MEMS sensor is fabricated

before during or after the CMOS circuit fabrication [36] Although these sensors offer

high-sensitivity and small parasitic capacitance they suffer from in-plane and out of

plane curling of the beam which reduces the capacitance between adjacent comb fingers

[37] In addition multiple processing steps are required for device fabrication

More than the other transduction methods MEMS capacitive displacement sensors are

attractive for practical implementation in an automobile and therefore is the selected

method for the vibration sensor presented in this work Their low temperature dependence

is an ideal characteristic for an automobile environment that faces varying temperature If

implemented with off-chip readout circuitry they require a standard foundry fabrication

process which provides a strong platform for mass production and low unit cost

3 SENSOR STRUCTURE

This chapter presents the physical structure of the three MEMS vibration sensors

which includes the displacement sensing capacitive comb drive and the micro-

flexureproof mass system First the capacitive comb drive and the three variations

implemented the transverse drive lateral drive and differential tri-plate drive are

discussed Next the mechanical structure is presented and includes the micro-flexure

selection quality factor and equations of motion Finally instability that places a limit

on the operating range of the sensor as a result of a gap closing capacitive comb drive is

discussed

31 SENSING

Capacitive sensors are based on the parallel plate capacitor and are commonly

implemented using inter-digitated comb fingers A capacitance can be realized if a voltage

is placed between two closely spaced plates If the fringing electric field is neglected the

capacitance of the parallel plate capacitor Cpp (F) is expressed as [28]

air pppp

k AC

dε

= (7)

where kair=1 is the dielectric constant of air at atmospheric pressure ε=88542times10-12 F m-1

is the permittivity of free space App is the overlapping area (m2) and d is the distance

between the two plates (m)

15

16

Figure 3-1 An inter-digitated comb drive where changes in capacitance are generated by either changes in gap distance x0 or in the overlap area y0timest

Figure 3-1

For a set of inter-digitated comb fingers or comb drive shown in the total

capacitance Ctotal is expressed as [28]

0 0

0 0

( ) ( )( ) (Total cy y y yC N t

)x x x xε

+ += + + minus

(8)

where N is the number of comb fingers on the movable comb y0 is the initial comb finger

overlap length (m) tc is the out of plane comb finger thickness (m) and x0 is the inital

finger gap (m) In this work the comb drive is implemented in three arrangements

transverse lateral and differential tri-plate In a transverse comb drive changes in

capacitance are generated by transverse movements along the x-axis resulting in changes

to the gap x0 between the movable and fixed comb fingers The nonlinear relation

between total capacitance CTotalT and displacement is expressed as

17

00 0

1 1( ) (Total TC N ty

)x x x xε

= + minus +

(9)

Although the drive sensitivity part is not constant an average sensitivity can be

assumed for small displacements [38]

Total TC xpart

In a lateral comb drive changes in capacitance are generated by lateral movements

along the y-axis resulting in changes to the overlapping area y0timestc The total capacitance

CTotalL between all comb fingers is expressed as

00

2 ( )Total LN tC yx

yε= +

(10)

The sensitivity of the lateral comb drive SL is expressed as

0

2Total LL

C N tSy x

εpart= =

part

(11)

The lateral comb drive was first introduced by W C Tang T H Nguyen and R T

Howe [33] in an electrostatic resonator and has since been used numerous times in both

drive and sense configurations It has the advantage of a fully linear response and

negligible instability but offers low sensitivity The sensitivity for the lateral drive L1

implemented in this work is 0038 pF microm-1

A differential tri-plate drive shown in Figure 3-2 which is a variation of the

transverse drive is implemented to improve sensitivity and reduce noise [39]

18

Figure 3-2 An inter-digitated comb drive in a differential tri-plate arrangement where x01 x02 and x03 are the gap distances and x01= x02ltlt x03 to improve sensitivity

In the tri-plate drive the at rest gap distances between the two capacitors formed by any

movable finger is not equivalent For a differential arrangement insuring that x01=

x02ltltx03 the linearity for small displacements (less than 1 microm) is enhanced over the case

of a single ended arrangement where x01=x02=x03=x0 which is equivalent to the

transverse drive This is shown graphically in Figure 3-3 where x01= x02=x03=2 microm in

the lower curve and x01= x02= 2 microm and x03= 15 microm in the upper curve A substantial

improvement in linearity is attained with the tri-plate drive

19

0

005

01

015

02

025

03

035

0 02 04 06 08 1

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-3 Comparison of the effect of gap distances on linearity for two case lsquorsquo represents case 1 where the fingers in are implemented in a differential tri-plate arrangement and x01= x02= 2 microm and x03= 15 microm lsquorsquo represents the case 2 where the fingers are implemented in a single ended arrangement and x01= x02=x03=2 microm

The tri-plate drive is also implemented in a differential arrangement where the output

is the change in capacitance rather than the total capacitance resulting from an external

excitation This reduces errors cause by off-axis excitation and common mode noise The

output of the differential tri-plate drive is expressed as [39]

Total Top Total BottomC C C∆ = minus where

003 01

1 1( ) (Total TopC N ty

)x x x xε

= +

minus +

(12)

20

002 03

1 1( ) (Total BottomC N ty

)x x x xε

= +

minus +

A graphic comparison of the capacitancedisplacement relation for sensors T1 and T2 is

shown in Figure 3-4 For these gap closing drives it is assumed that for displacements

less that 1 microm the relation is linear indicated by the dashed line above each curve The

sensitivity within this region is approximately 0046 pF microm-1 and 0253 pF microm-1 for T1

and T2 respectively The order of magnitude improvement is attributed to the differential

tri-plate arrangement and thicker comb finger

0

02

04

06

08

1

12

14

16

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-4 lsquorsquo represents the output of the differential tri-plate drive while lsquorsquo represents the output of the transverse drive The dashed line represents the linear approximation for displacements less that 1 microm For the purpose of comparison the output of the transverse drive was equalized to zero at zero displacement

21

In this work the transverse drive is first implemented in sensor T1 using the

PolyMUMPS fabrication process Then in an effort to linearize the sensor response the

lateral drive is used in L1 using SoiMUMPS The reduction in drive sensitivity due to the

lateral drive is further reduced by increasing the gap distance to 3 microm which is

recommended by the foundry to ensure a successful device These reductions are

balanced by increasing the number of comb fingers N and selecting a substrate thickness

that is 125 times large than T1 The third device T2 is fabricated using SoiMUMPS and

uses the tri-plate differential drive This design capitalizes on the increase in sensitivity of

a gap closing drive and larger substrate thickness attaining an order of magnitude

increase in sensitivity over T1 and L1 The specifications for each drive along with the

successive improvement in drive sensitivity are outlined in Table 3-1

Table 3-1 Specifications for each capacitive comb drive implemented in this work T1 L1 and T2

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Comb Finger Thickness tc (microm) 2 25 25 Gap Distance x0 (microm) 2 3 2 Number of Fingers N 200 256 30 Overlap length y0 (microm) 45 10 70 Drive Sensitivity (pF microm-1) 0046 0038 0273

32 MECHANICAL ANALYSIS

321 MICRO-FLEXURE SELECTION

The sensor structure consists of a proof mass connected to a compliant micro-flexure

that itself is anchored to the sensor substrate The proof mass displaces in response to an

22

applied acceleration and this displacement is sensed using the capacitive comb drive

attached to the proof mass

Proof mass

Folded beam micro-flexure

Anchor

Figure 3-5 Sensor Structure

This section focuses on micro-flexure selection as it should exhibit a large linear

displacement range and have a high lateral stiffness ratio to reduce cross-axis sensitivity

A number of micro-flexures have been used in MEMS research these include the

clamped-clamped crab-leg and folded beam shown in Figure 3-6-through 3-7

The clamped-clamped flexure exhibits a high stiffness ratio however for large lateral

displacements extensional axial forces buildup within the beams resulting in a non-linear

force-displacement relation [40] Within the linear regime the stiffness of the clamped-

clamped flexure is expressed as [40]

3

3b

yEt wk

L=

(13)

23

where E is the Youngrsquos Modulus of the beam material (N m-2) tb is the beam out of

plane thickness (m) w is the beam width (m) and L is the beam length (m) The stiffness

ratio is expressed as [40]

2x

y

k Lk w

=

(14)

If L is two orders of magnitude greater than w this results in a ratio of approximately

10000 The linear displacement range is determined using Finite Element Analysis (FEA)

of a clamped-clamped flexure with L=387 microm w=4 microm and tb=2 microm The result of this

simulation is plotted in and compared to a structure with constant stiffness ky

= 035 N m-1 predicted by Eq (13) The linear relation is valid for small deflections that

are less than 2 microm In addition this micro-flexure is susceptible to buckling as a result of

a compressive residual stress that may arise due to the fabrication process [41]

Figure 3-9

In the crab leg design the thigh section is meant to relieve the extensional axial

stress postponing the occurrence of non-linearity until greater lateral displacement [42]

The stiffness of the crab leg flexure is expressed as [43]

3

44

b s s ty

s s t

Et w L LkL L L

κκ

+= +

(15)

where κ=IsIt=(wswt)3 is a ratio of the second moment of inertia of the shin (s) and thigh

(t) ws and wt represent the width and Ls and Lt represent the length of the shin and thigh

regions respectively The stiffness ratio is expressed as [43]

34

4t s s tx

y s t s

w L L Lkk w L L L

αα t

+= +

(16)

The linear displacement range of a crab leg flexure with Ls=285 microm Lt=30 microm ws=wt=3

microm and tb=2 microm is determined using FEA The results are plotted in Figure 3-9 and

compared to a structure of constant stiffness ky=035 N m-1 predicted by Eq (15) The

24

non-linearity is greatly reduced when compared to the clamp-clamped flexure with Eq

accurately predicting the deflection until 4 microm However the flexure is more

susceptible to off-axis deflection as there is an order of magnitude reduction in the

stiffness ratio as compared to the clamped-clamped flexure

(15)

The folded beam flexure stands out as it is designed to exhibit a large linear

displacement range and to relieve residual stress due to the manufacturing process by

allowing the beams to expand and contract along the axial direction [40] It was first used

in an electrostatic resonator and has subsequently been used for electrostatic actuation in

a magnetic hard drives and micro-positioning [3344] If the truss section is rigid

ItrussgtgtIbeam then the stiffness of the flexure is expressed as [4043]

3

32b

yEt wk

L=

(17)

The stiffness ratio is expressed as [40]

2x

y

k Lk w

=

(18)

Figure 3-6 Clamped-clamped flexure

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

3 SENSOR STRUCTURE

This chapter presents the physical structure of the three MEMS vibration sensors

which includes the displacement sensing capacitive comb drive and the micro-

flexureproof mass system First the capacitive comb drive and the three variations

implemented the transverse drive lateral drive and differential tri-plate drive are

discussed Next the mechanical structure is presented and includes the micro-flexure

selection quality factor and equations of motion Finally instability that places a limit

on the operating range of the sensor as a result of a gap closing capacitive comb drive is

discussed

31 SENSING

Capacitive sensors are based on the parallel plate capacitor and are commonly

implemented using inter-digitated comb fingers A capacitance can be realized if a voltage

is placed between two closely spaced plates If the fringing electric field is neglected the

capacitance of the parallel plate capacitor Cpp (F) is expressed as [28]

air pppp

k AC

dε

= (7)

where kair=1 is the dielectric constant of air at atmospheric pressure ε=88542times10-12 F m-1

is the permittivity of free space App is the overlapping area (m2) and d is the distance

between the two plates (m)

15

16

Figure 3-1 An inter-digitated comb drive where changes in capacitance are generated by either changes in gap distance x0 or in the overlap area y0timest

Figure 3-1

For a set of inter-digitated comb fingers or comb drive shown in the total

capacitance Ctotal is expressed as [28]

0 0

0 0

( ) ( )( ) (Total cy y y yC N t

)x x x xε

+ += + + minus

(8)

where N is the number of comb fingers on the movable comb y0 is the initial comb finger

overlap length (m) tc is the out of plane comb finger thickness (m) and x0 is the inital

finger gap (m) In this work the comb drive is implemented in three arrangements

transverse lateral and differential tri-plate In a transverse comb drive changes in

capacitance are generated by transverse movements along the x-axis resulting in changes

to the gap x0 between the movable and fixed comb fingers The nonlinear relation

between total capacitance CTotalT and displacement is expressed as

17

00 0

1 1( ) (Total TC N ty

)x x x xε

= + minus +

(9)

Although the drive sensitivity part is not constant an average sensitivity can be

assumed for small displacements [38]

Total TC xpart

In a lateral comb drive changes in capacitance are generated by lateral movements

along the y-axis resulting in changes to the overlapping area y0timestc The total capacitance

CTotalL between all comb fingers is expressed as

00

2 ( )Total LN tC yx

yε= +

(10)

The sensitivity of the lateral comb drive SL is expressed as

0

2Total LL

C N tSy x

εpart= =

part

(11)

The lateral comb drive was first introduced by W C Tang T H Nguyen and R T

Howe [33] in an electrostatic resonator and has since been used numerous times in both

drive and sense configurations It has the advantage of a fully linear response and

negligible instability but offers low sensitivity The sensitivity for the lateral drive L1

implemented in this work is 0038 pF microm-1

A differential tri-plate drive shown in Figure 3-2 which is a variation of the

transverse drive is implemented to improve sensitivity and reduce noise [39]

18

Figure 3-2 An inter-digitated comb drive in a differential tri-plate arrangement where x01 x02 and x03 are the gap distances and x01= x02ltlt x03 to improve sensitivity

In the tri-plate drive the at rest gap distances between the two capacitors formed by any

movable finger is not equivalent For a differential arrangement insuring that x01=

x02ltltx03 the linearity for small displacements (less than 1 microm) is enhanced over the case

of a single ended arrangement where x01=x02=x03=x0 which is equivalent to the

transverse drive This is shown graphically in Figure 3-3 where x01= x02=x03=2 microm in

the lower curve and x01= x02= 2 microm and x03= 15 microm in the upper curve A substantial

improvement in linearity is attained with the tri-plate drive

19

0

005

01

015

02

025

03

035

0 02 04 06 08 1

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-3 Comparison of the effect of gap distances on linearity for two case lsquorsquo represents case 1 where the fingers in are implemented in a differential tri-plate arrangement and x01= x02= 2 microm and x03= 15 microm lsquorsquo represents the case 2 where the fingers are implemented in a single ended arrangement and x01= x02=x03=2 microm

The tri-plate drive is also implemented in a differential arrangement where the output

is the change in capacitance rather than the total capacitance resulting from an external

excitation This reduces errors cause by off-axis excitation and common mode noise The

output of the differential tri-plate drive is expressed as [39]

Total Top Total BottomC C C∆ = minus where

003 01

1 1( ) (Total TopC N ty

)x x x xε

= +

minus +

(12)

20

002 03

1 1( ) (Total BottomC N ty

)x x x xε

= +

minus +

A graphic comparison of the capacitancedisplacement relation for sensors T1 and T2 is

shown in Figure 3-4 For these gap closing drives it is assumed that for displacements

less that 1 microm the relation is linear indicated by the dashed line above each curve The

sensitivity within this region is approximately 0046 pF microm-1 and 0253 pF microm-1 for T1

and T2 respectively The order of magnitude improvement is attributed to the differential

tri-plate arrangement and thicker comb finger

0

02

04

06

08

1

12

14

16

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-4 lsquorsquo represents the output of the differential tri-plate drive while lsquorsquo represents the output of the transverse drive The dashed line represents the linear approximation for displacements less that 1 microm For the purpose of comparison the output of the transverse drive was equalized to zero at zero displacement

21

In this work the transverse drive is first implemented in sensor T1 using the

PolyMUMPS fabrication process Then in an effort to linearize the sensor response the

lateral drive is used in L1 using SoiMUMPS The reduction in drive sensitivity due to the

lateral drive is further reduced by increasing the gap distance to 3 microm which is

recommended by the foundry to ensure a successful device These reductions are

balanced by increasing the number of comb fingers N and selecting a substrate thickness

that is 125 times large than T1 The third device T2 is fabricated using SoiMUMPS and

uses the tri-plate differential drive This design capitalizes on the increase in sensitivity of

a gap closing drive and larger substrate thickness attaining an order of magnitude

increase in sensitivity over T1 and L1 The specifications for each drive along with the

successive improvement in drive sensitivity are outlined in Table 3-1

Table 3-1 Specifications for each capacitive comb drive implemented in this work T1 L1 and T2

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Comb Finger Thickness tc (microm) 2 25 25 Gap Distance x0 (microm) 2 3 2 Number of Fingers N 200 256 30 Overlap length y0 (microm) 45 10 70 Drive Sensitivity (pF microm-1) 0046 0038 0273

32 MECHANICAL ANALYSIS

321 MICRO-FLEXURE SELECTION

The sensor structure consists of a proof mass connected to a compliant micro-flexure

that itself is anchored to the sensor substrate The proof mass displaces in response to an

22

applied acceleration and this displacement is sensed using the capacitive comb drive

attached to the proof mass

Proof mass

Folded beam micro-flexure

Anchor

Figure 3-5 Sensor Structure

This section focuses on micro-flexure selection as it should exhibit a large linear

displacement range and have a high lateral stiffness ratio to reduce cross-axis sensitivity

A number of micro-flexures have been used in MEMS research these include the

clamped-clamped crab-leg and folded beam shown in Figure 3-6-through 3-7

The clamped-clamped flexure exhibits a high stiffness ratio however for large lateral

displacements extensional axial forces buildup within the beams resulting in a non-linear

force-displacement relation [40] Within the linear regime the stiffness of the clamped-

clamped flexure is expressed as [40]

3

3b

yEt wk

L=

(13)

23

where E is the Youngrsquos Modulus of the beam material (N m-2) tb is the beam out of

plane thickness (m) w is the beam width (m) and L is the beam length (m) The stiffness

ratio is expressed as [40]

2x

y

k Lk w

=

(14)

If L is two orders of magnitude greater than w this results in a ratio of approximately

10000 The linear displacement range is determined using Finite Element Analysis (FEA)

of a clamped-clamped flexure with L=387 microm w=4 microm and tb=2 microm The result of this

simulation is plotted in and compared to a structure with constant stiffness ky

= 035 N m-1 predicted by Eq (13) The linear relation is valid for small deflections that

are less than 2 microm In addition this micro-flexure is susceptible to buckling as a result of

a compressive residual stress that may arise due to the fabrication process [41]

Figure 3-9

In the crab leg design the thigh section is meant to relieve the extensional axial

stress postponing the occurrence of non-linearity until greater lateral displacement [42]

The stiffness of the crab leg flexure is expressed as [43]

3

44

b s s ty

s s t

Et w L LkL L L

κκ

+= +

(15)

where κ=IsIt=(wswt)3 is a ratio of the second moment of inertia of the shin (s) and thigh

(t) ws and wt represent the width and Ls and Lt represent the length of the shin and thigh

regions respectively The stiffness ratio is expressed as [43]

34

4t s s tx

y s t s

w L L Lkk w L L L

αα t

+= +

(16)

The linear displacement range of a crab leg flexure with Ls=285 microm Lt=30 microm ws=wt=3

microm and tb=2 microm is determined using FEA The results are plotted in Figure 3-9 and

compared to a structure of constant stiffness ky=035 N m-1 predicted by Eq (15) The

24

non-linearity is greatly reduced when compared to the clamp-clamped flexure with Eq

accurately predicting the deflection until 4 microm However the flexure is more

susceptible to off-axis deflection as there is an order of magnitude reduction in the

stiffness ratio as compared to the clamped-clamped flexure

(15)

The folded beam flexure stands out as it is designed to exhibit a large linear

displacement range and to relieve residual stress due to the manufacturing process by

allowing the beams to expand and contract along the axial direction [40] It was first used

in an electrostatic resonator and has subsequently been used for electrostatic actuation in

a magnetic hard drives and micro-positioning [3344] If the truss section is rigid

ItrussgtgtIbeam then the stiffness of the flexure is expressed as [4043]

3

32b

yEt wk

L=

(17)

The stiffness ratio is expressed as [40]

2x

y

k Lk w

=

(18)

Figure 3-6 Clamped-clamped flexure

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

16

Figure 3-1 An inter-digitated comb drive where changes in capacitance are generated by either changes in gap distance x0 or in the overlap area y0timest

Figure 3-1

For a set of inter-digitated comb fingers or comb drive shown in the total

capacitance Ctotal is expressed as [28]

0 0

0 0

( ) ( )( ) (Total cy y y yC N t

)x x x xε

+ += + + minus

(8)

where N is the number of comb fingers on the movable comb y0 is the initial comb finger

overlap length (m) tc is the out of plane comb finger thickness (m) and x0 is the inital

finger gap (m) In this work the comb drive is implemented in three arrangements

transverse lateral and differential tri-plate In a transverse comb drive changes in

capacitance are generated by transverse movements along the x-axis resulting in changes

to the gap x0 between the movable and fixed comb fingers The nonlinear relation

between total capacitance CTotalT and displacement is expressed as

17

00 0

1 1( ) (Total TC N ty

)x x x xε

= + minus +

(9)

Although the drive sensitivity part is not constant an average sensitivity can be

assumed for small displacements [38]

Total TC xpart

In a lateral comb drive changes in capacitance are generated by lateral movements

along the y-axis resulting in changes to the overlapping area y0timestc The total capacitance

CTotalL between all comb fingers is expressed as

00

2 ( )Total LN tC yx

yε= +

(10)

The sensitivity of the lateral comb drive SL is expressed as

0

2Total LL

C N tSy x

εpart= =

part

(11)

The lateral comb drive was first introduced by W C Tang T H Nguyen and R T

Howe [33] in an electrostatic resonator and has since been used numerous times in both

drive and sense configurations It has the advantage of a fully linear response and

negligible instability but offers low sensitivity The sensitivity for the lateral drive L1

implemented in this work is 0038 pF microm-1

A differential tri-plate drive shown in Figure 3-2 which is a variation of the

transverse drive is implemented to improve sensitivity and reduce noise [39]

18

Figure 3-2 An inter-digitated comb drive in a differential tri-plate arrangement where x01 x02 and x03 are the gap distances and x01= x02ltlt x03 to improve sensitivity

In the tri-plate drive the at rest gap distances between the two capacitors formed by any

movable finger is not equivalent For a differential arrangement insuring that x01=

x02ltltx03 the linearity for small displacements (less than 1 microm) is enhanced over the case

of a single ended arrangement where x01=x02=x03=x0 which is equivalent to the

transverse drive This is shown graphically in Figure 3-3 where x01= x02=x03=2 microm in

the lower curve and x01= x02= 2 microm and x03= 15 microm in the upper curve A substantial

improvement in linearity is attained with the tri-plate drive

19

0

005

01

015

02

025

03

035

0 02 04 06 08 1

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-3 Comparison of the effect of gap distances on linearity for two case lsquorsquo represents case 1 where the fingers in are implemented in a differential tri-plate arrangement and x01= x02= 2 microm and x03= 15 microm lsquorsquo represents the case 2 where the fingers are implemented in a single ended arrangement and x01= x02=x03=2 microm

The tri-plate drive is also implemented in a differential arrangement where the output

is the change in capacitance rather than the total capacitance resulting from an external

excitation This reduces errors cause by off-axis excitation and common mode noise The

output of the differential tri-plate drive is expressed as [39]

Total Top Total BottomC C C∆ = minus where

003 01

1 1( ) (Total TopC N ty

)x x x xε

= +

minus +

(12)

20

002 03

1 1( ) (Total BottomC N ty

)x x x xε

= +

minus +

A graphic comparison of the capacitancedisplacement relation for sensors T1 and T2 is

shown in Figure 3-4 For these gap closing drives it is assumed that for displacements

less that 1 microm the relation is linear indicated by the dashed line above each curve The

sensitivity within this region is approximately 0046 pF microm-1 and 0253 pF microm-1 for T1

and T2 respectively The order of magnitude improvement is attributed to the differential

tri-plate arrangement and thicker comb finger

0

02

04

06

08

1

12

14

16

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-4 lsquorsquo represents the output of the differential tri-plate drive while lsquorsquo represents the output of the transverse drive The dashed line represents the linear approximation for displacements less that 1 microm For the purpose of comparison the output of the transverse drive was equalized to zero at zero displacement

21

In this work the transverse drive is first implemented in sensor T1 using the

PolyMUMPS fabrication process Then in an effort to linearize the sensor response the

lateral drive is used in L1 using SoiMUMPS The reduction in drive sensitivity due to the

lateral drive is further reduced by increasing the gap distance to 3 microm which is

recommended by the foundry to ensure a successful device These reductions are

balanced by increasing the number of comb fingers N and selecting a substrate thickness

that is 125 times large than T1 The third device T2 is fabricated using SoiMUMPS and

uses the tri-plate differential drive This design capitalizes on the increase in sensitivity of

a gap closing drive and larger substrate thickness attaining an order of magnitude

increase in sensitivity over T1 and L1 The specifications for each drive along with the

successive improvement in drive sensitivity are outlined in Table 3-1

Table 3-1 Specifications for each capacitive comb drive implemented in this work T1 L1 and T2

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Comb Finger Thickness tc (microm) 2 25 25 Gap Distance x0 (microm) 2 3 2 Number of Fingers N 200 256 30 Overlap length y0 (microm) 45 10 70 Drive Sensitivity (pF microm-1) 0046 0038 0273

32 MECHANICAL ANALYSIS

321 MICRO-FLEXURE SELECTION

The sensor structure consists of a proof mass connected to a compliant micro-flexure

that itself is anchored to the sensor substrate The proof mass displaces in response to an

22

applied acceleration and this displacement is sensed using the capacitive comb drive

attached to the proof mass

Proof mass

Folded beam micro-flexure

Anchor

Figure 3-5 Sensor Structure

This section focuses on micro-flexure selection as it should exhibit a large linear

displacement range and have a high lateral stiffness ratio to reduce cross-axis sensitivity

A number of micro-flexures have been used in MEMS research these include the

clamped-clamped crab-leg and folded beam shown in Figure 3-6-through 3-7

The clamped-clamped flexure exhibits a high stiffness ratio however for large lateral

displacements extensional axial forces buildup within the beams resulting in a non-linear

force-displacement relation [40] Within the linear regime the stiffness of the clamped-

clamped flexure is expressed as [40]

3

3b

yEt wk

L=

(13)

23

where E is the Youngrsquos Modulus of the beam material (N m-2) tb is the beam out of

plane thickness (m) w is the beam width (m) and L is the beam length (m) The stiffness

ratio is expressed as [40]

2x

y

k Lk w

=

(14)

If L is two orders of magnitude greater than w this results in a ratio of approximately

10000 The linear displacement range is determined using Finite Element Analysis (FEA)

of a clamped-clamped flexure with L=387 microm w=4 microm and tb=2 microm The result of this

simulation is plotted in and compared to a structure with constant stiffness ky

= 035 N m-1 predicted by Eq (13) The linear relation is valid for small deflections that

are less than 2 microm In addition this micro-flexure is susceptible to buckling as a result of

a compressive residual stress that may arise due to the fabrication process [41]

Figure 3-9

In the crab leg design the thigh section is meant to relieve the extensional axial

stress postponing the occurrence of non-linearity until greater lateral displacement [42]

The stiffness of the crab leg flexure is expressed as [43]

3

44

b s s ty

s s t

Et w L LkL L L

κκ

+= +

(15)

where κ=IsIt=(wswt)3 is a ratio of the second moment of inertia of the shin (s) and thigh

(t) ws and wt represent the width and Ls and Lt represent the length of the shin and thigh

regions respectively The stiffness ratio is expressed as [43]

34

4t s s tx

y s t s

w L L Lkk w L L L

αα t

+= +

(16)

The linear displacement range of a crab leg flexure with Ls=285 microm Lt=30 microm ws=wt=3

microm and tb=2 microm is determined using FEA The results are plotted in Figure 3-9 and

compared to a structure of constant stiffness ky=035 N m-1 predicted by Eq (15) The

24

non-linearity is greatly reduced when compared to the clamp-clamped flexure with Eq

accurately predicting the deflection until 4 microm However the flexure is more

susceptible to off-axis deflection as there is an order of magnitude reduction in the

stiffness ratio as compared to the clamped-clamped flexure

(15)

The folded beam flexure stands out as it is designed to exhibit a large linear

displacement range and to relieve residual stress due to the manufacturing process by

allowing the beams to expand and contract along the axial direction [40] It was first used

in an electrostatic resonator and has subsequently been used for electrostatic actuation in

a magnetic hard drives and micro-positioning [3344] If the truss section is rigid

ItrussgtgtIbeam then the stiffness of the flexure is expressed as [4043]

3

32b

yEt wk

L=

(17)

The stiffness ratio is expressed as [40]

2x

y

k Lk w

=

(18)

Figure 3-6 Clamped-clamped flexure

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

17

00 0

1 1( ) (Total TC N ty

)x x x xε

= + minus +

(9)

Although the drive sensitivity part is not constant an average sensitivity can be

assumed for small displacements [38]

Total TC xpart

In a lateral comb drive changes in capacitance are generated by lateral movements

along the y-axis resulting in changes to the overlapping area y0timestc The total capacitance

CTotalL between all comb fingers is expressed as

00

2 ( )Total LN tC yx

yε= +

(10)

The sensitivity of the lateral comb drive SL is expressed as

0

2Total LL

C N tSy x

εpart= =

part

(11)

The lateral comb drive was first introduced by W C Tang T H Nguyen and R T

Howe [33] in an electrostatic resonator and has since been used numerous times in both

drive and sense configurations It has the advantage of a fully linear response and

negligible instability but offers low sensitivity The sensitivity for the lateral drive L1

implemented in this work is 0038 pF microm-1

A differential tri-plate drive shown in Figure 3-2 which is a variation of the

transverse drive is implemented to improve sensitivity and reduce noise [39]

18

Figure 3-2 An inter-digitated comb drive in a differential tri-plate arrangement where x01 x02 and x03 are the gap distances and x01= x02ltlt x03 to improve sensitivity

In the tri-plate drive the at rest gap distances between the two capacitors formed by any

movable finger is not equivalent For a differential arrangement insuring that x01=

x02ltltx03 the linearity for small displacements (less than 1 microm) is enhanced over the case

of a single ended arrangement where x01=x02=x03=x0 which is equivalent to the

transverse drive This is shown graphically in Figure 3-3 where x01= x02=x03=2 microm in

the lower curve and x01= x02= 2 microm and x03= 15 microm in the upper curve A substantial

improvement in linearity is attained with the tri-plate drive

19

0

005

01

015

02

025

03

035

0 02 04 06 08 1

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-3 Comparison of the effect of gap distances on linearity for two case lsquorsquo represents case 1 where the fingers in are implemented in a differential tri-plate arrangement and x01= x02= 2 microm and x03= 15 microm lsquorsquo represents the case 2 where the fingers are implemented in a single ended arrangement and x01= x02=x03=2 microm

The tri-plate drive is also implemented in a differential arrangement where the output

is the change in capacitance rather than the total capacitance resulting from an external

excitation This reduces errors cause by off-axis excitation and common mode noise The

output of the differential tri-plate drive is expressed as [39]

Total Top Total BottomC C C∆ = minus where

003 01

1 1( ) (Total TopC N ty

)x x x xε

= +

minus +

(12)

20

002 03

1 1( ) (Total BottomC N ty

)x x x xε

= +

minus +

A graphic comparison of the capacitancedisplacement relation for sensors T1 and T2 is

shown in Figure 3-4 For these gap closing drives it is assumed that for displacements

less that 1 microm the relation is linear indicated by the dashed line above each curve The

sensitivity within this region is approximately 0046 pF microm-1 and 0253 pF microm-1 for T1

and T2 respectively The order of magnitude improvement is attributed to the differential

tri-plate arrangement and thicker comb finger

0

02

04

06

08

1

12

14

16

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-4 lsquorsquo represents the output of the differential tri-plate drive while lsquorsquo represents the output of the transverse drive The dashed line represents the linear approximation for displacements less that 1 microm For the purpose of comparison the output of the transverse drive was equalized to zero at zero displacement

21

In this work the transverse drive is first implemented in sensor T1 using the

PolyMUMPS fabrication process Then in an effort to linearize the sensor response the

lateral drive is used in L1 using SoiMUMPS The reduction in drive sensitivity due to the

lateral drive is further reduced by increasing the gap distance to 3 microm which is

recommended by the foundry to ensure a successful device These reductions are

balanced by increasing the number of comb fingers N and selecting a substrate thickness

that is 125 times large than T1 The third device T2 is fabricated using SoiMUMPS and

uses the tri-plate differential drive This design capitalizes on the increase in sensitivity of

a gap closing drive and larger substrate thickness attaining an order of magnitude

increase in sensitivity over T1 and L1 The specifications for each drive along with the

successive improvement in drive sensitivity are outlined in Table 3-1

Table 3-1 Specifications for each capacitive comb drive implemented in this work T1 L1 and T2

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Comb Finger Thickness tc (microm) 2 25 25 Gap Distance x0 (microm) 2 3 2 Number of Fingers N 200 256 30 Overlap length y0 (microm) 45 10 70 Drive Sensitivity (pF microm-1) 0046 0038 0273

32 MECHANICAL ANALYSIS

321 MICRO-FLEXURE SELECTION

The sensor structure consists of a proof mass connected to a compliant micro-flexure

that itself is anchored to the sensor substrate The proof mass displaces in response to an

22

applied acceleration and this displacement is sensed using the capacitive comb drive

attached to the proof mass

Proof mass

Folded beam micro-flexure

Anchor

Figure 3-5 Sensor Structure

This section focuses on micro-flexure selection as it should exhibit a large linear

displacement range and have a high lateral stiffness ratio to reduce cross-axis sensitivity

A number of micro-flexures have been used in MEMS research these include the

clamped-clamped crab-leg and folded beam shown in Figure 3-6-through 3-7

The clamped-clamped flexure exhibits a high stiffness ratio however for large lateral

displacements extensional axial forces buildup within the beams resulting in a non-linear

force-displacement relation [40] Within the linear regime the stiffness of the clamped-

clamped flexure is expressed as [40]

3

3b

yEt wk

L=

(13)

23

where E is the Youngrsquos Modulus of the beam material (N m-2) tb is the beam out of

plane thickness (m) w is the beam width (m) and L is the beam length (m) The stiffness

ratio is expressed as [40]

2x

y

k Lk w

=

(14)

If L is two orders of magnitude greater than w this results in a ratio of approximately

10000 The linear displacement range is determined using Finite Element Analysis (FEA)

of a clamped-clamped flexure with L=387 microm w=4 microm and tb=2 microm The result of this

simulation is plotted in and compared to a structure with constant stiffness ky

= 035 N m-1 predicted by Eq (13) The linear relation is valid for small deflections that

are less than 2 microm In addition this micro-flexure is susceptible to buckling as a result of

a compressive residual stress that may arise due to the fabrication process [41]

Figure 3-9

In the crab leg design the thigh section is meant to relieve the extensional axial

stress postponing the occurrence of non-linearity until greater lateral displacement [42]

The stiffness of the crab leg flexure is expressed as [43]

3

44

b s s ty

s s t

Et w L LkL L L

κκ

+= +

(15)

where κ=IsIt=(wswt)3 is a ratio of the second moment of inertia of the shin (s) and thigh

(t) ws and wt represent the width and Ls and Lt represent the length of the shin and thigh

regions respectively The stiffness ratio is expressed as [43]

34

4t s s tx

y s t s

w L L Lkk w L L L

αα t

+= +

(16)

The linear displacement range of a crab leg flexure with Ls=285 microm Lt=30 microm ws=wt=3

microm and tb=2 microm is determined using FEA The results are plotted in Figure 3-9 and

compared to a structure of constant stiffness ky=035 N m-1 predicted by Eq (15) The

24

non-linearity is greatly reduced when compared to the clamp-clamped flexure with Eq

accurately predicting the deflection until 4 microm However the flexure is more

susceptible to off-axis deflection as there is an order of magnitude reduction in the

stiffness ratio as compared to the clamped-clamped flexure

(15)

The folded beam flexure stands out as it is designed to exhibit a large linear

displacement range and to relieve residual stress due to the manufacturing process by

allowing the beams to expand and contract along the axial direction [40] It was first used

in an electrostatic resonator and has subsequently been used for electrostatic actuation in

a magnetic hard drives and micro-positioning [3344] If the truss section is rigid

ItrussgtgtIbeam then the stiffness of the flexure is expressed as [4043]

3

32b

yEt wk

L=

(17)

The stiffness ratio is expressed as [40]

2x

y

k Lk w

=

(18)

Figure 3-6 Clamped-clamped flexure

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

18

Figure 3-2 An inter-digitated comb drive in a differential tri-plate arrangement where x01 x02 and x03 are the gap distances and x01= x02ltlt x03 to improve sensitivity

In the tri-plate drive the at rest gap distances between the two capacitors formed by any

movable finger is not equivalent For a differential arrangement insuring that x01=

x02ltltx03 the linearity for small displacements (less than 1 microm) is enhanced over the case

of a single ended arrangement where x01=x02=x03=x0 which is equivalent to the

transverse drive This is shown graphically in Figure 3-3 where x01= x02=x03=2 microm in

the lower curve and x01= x02= 2 microm and x03= 15 microm in the upper curve A substantial

improvement in linearity is attained with the tri-plate drive

19

0

005

01

015

02

025

03

035

0 02 04 06 08 1

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-3 Comparison of the effect of gap distances on linearity for two case lsquorsquo represents case 1 where the fingers in are implemented in a differential tri-plate arrangement and x01= x02= 2 microm and x03= 15 microm lsquorsquo represents the case 2 where the fingers are implemented in a single ended arrangement and x01= x02=x03=2 microm

The tri-plate drive is also implemented in a differential arrangement where the output

is the change in capacitance rather than the total capacitance resulting from an external

excitation This reduces errors cause by off-axis excitation and common mode noise The

output of the differential tri-plate drive is expressed as [39]

Total Top Total BottomC C C∆ = minus where

003 01

1 1( ) (Total TopC N ty

)x x x xε

= +

minus +

(12)

20

002 03

1 1( ) (Total BottomC N ty

)x x x xε

= +

minus +

A graphic comparison of the capacitancedisplacement relation for sensors T1 and T2 is

shown in Figure 3-4 For these gap closing drives it is assumed that for displacements

less that 1 microm the relation is linear indicated by the dashed line above each curve The

sensitivity within this region is approximately 0046 pF microm-1 and 0253 pF microm-1 for T1

and T2 respectively The order of magnitude improvement is attributed to the differential

tri-plate arrangement and thicker comb finger

0

02

04

06

08

1

12

14

16

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-4 lsquorsquo represents the output of the differential tri-plate drive while lsquorsquo represents the output of the transverse drive The dashed line represents the linear approximation for displacements less that 1 microm For the purpose of comparison the output of the transverse drive was equalized to zero at zero displacement

21

In this work the transverse drive is first implemented in sensor T1 using the

PolyMUMPS fabrication process Then in an effort to linearize the sensor response the

lateral drive is used in L1 using SoiMUMPS The reduction in drive sensitivity due to the

lateral drive is further reduced by increasing the gap distance to 3 microm which is

recommended by the foundry to ensure a successful device These reductions are

balanced by increasing the number of comb fingers N and selecting a substrate thickness

that is 125 times large than T1 The third device T2 is fabricated using SoiMUMPS and

uses the tri-plate differential drive This design capitalizes on the increase in sensitivity of

a gap closing drive and larger substrate thickness attaining an order of magnitude

increase in sensitivity over T1 and L1 The specifications for each drive along with the

successive improvement in drive sensitivity are outlined in Table 3-1

Table 3-1 Specifications for each capacitive comb drive implemented in this work T1 L1 and T2

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Comb Finger Thickness tc (microm) 2 25 25 Gap Distance x0 (microm) 2 3 2 Number of Fingers N 200 256 30 Overlap length y0 (microm) 45 10 70 Drive Sensitivity (pF microm-1) 0046 0038 0273

32 MECHANICAL ANALYSIS

321 MICRO-FLEXURE SELECTION

The sensor structure consists of a proof mass connected to a compliant micro-flexure

that itself is anchored to the sensor substrate The proof mass displaces in response to an

22

applied acceleration and this displacement is sensed using the capacitive comb drive

attached to the proof mass

Proof mass

Folded beam micro-flexure

Anchor

Figure 3-5 Sensor Structure

This section focuses on micro-flexure selection as it should exhibit a large linear

displacement range and have a high lateral stiffness ratio to reduce cross-axis sensitivity

A number of micro-flexures have been used in MEMS research these include the

clamped-clamped crab-leg and folded beam shown in Figure 3-6-through 3-7

The clamped-clamped flexure exhibits a high stiffness ratio however for large lateral

displacements extensional axial forces buildup within the beams resulting in a non-linear

force-displacement relation [40] Within the linear regime the stiffness of the clamped-

clamped flexure is expressed as [40]

3

3b

yEt wk

L=

(13)

23

where E is the Youngrsquos Modulus of the beam material (N m-2) tb is the beam out of

plane thickness (m) w is the beam width (m) and L is the beam length (m) The stiffness

ratio is expressed as [40]

2x

y

k Lk w

=

(14)

If L is two orders of magnitude greater than w this results in a ratio of approximately

10000 The linear displacement range is determined using Finite Element Analysis (FEA)

of a clamped-clamped flexure with L=387 microm w=4 microm and tb=2 microm The result of this

simulation is plotted in and compared to a structure with constant stiffness ky

= 035 N m-1 predicted by Eq (13) The linear relation is valid for small deflections that

are less than 2 microm In addition this micro-flexure is susceptible to buckling as a result of

a compressive residual stress that may arise due to the fabrication process [41]

Figure 3-9

In the crab leg design the thigh section is meant to relieve the extensional axial

stress postponing the occurrence of non-linearity until greater lateral displacement [42]

The stiffness of the crab leg flexure is expressed as [43]

3

44

b s s ty

s s t

Et w L LkL L L

κκ

+= +

(15)

where κ=IsIt=(wswt)3 is a ratio of the second moment of inertia of the shin (s) and thigh

(t) ws and wt represent the width and Ls and Lt represent the length of the shin and thigh

regions respectively The stiffness ratio is expressed as [43]

34

4t s s tx

y s t s

w L L Lkk w L L L

αα t

+= +

(16)

The linear displacement range of a crab leg flexure with Ls=285 microm Lt=30 microm ws=wt=3

microm and tb=2 microm is determined using FEA The results are plotted in Figure 3-9 and

compared to a structure of constant stiffness ky=035 N m-1 predicted by Eq (15) The

24

non-linearity is greatly reduced when compared to the clamp-clamped flexure with Eq

accurately predicting the deflection until 4 microm However the flexure is more

susceptible to off-axis deflection as there is an order of magnitude reduction in the

stiffness ratio as compared to the clamped-clamped flexure

(15)

The folded beam flexure stands out as it is designed to exhibit a large linear

displacement range and to relieve residual stress due to the manufacturing process by

allowing the beams to expand and contract along the axial direction [40] It was first used

in an electrostatic resonator and has subsequently been used for electrostatic actuation in

a magnetic hard drives and micro-positioning [3344] If the truss section is rigid

ItrussgtgtIbeam then the stiffness of the flexure is expressed as [4043]

3

32b

yEt wk

L=

(17)

The stiffness ratio is expressed as [40]

2x

y

k Lk w

=

(18)

Figure 3-6 Clamped-clamped flexure

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

19

0

005

01

015

02

025

03

035

0 02 04 06 08 1

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-3 Comparison of the effect of gap distances on linearity for two case lsquorsquo represents case 1 where the fingers in are implemented in a differential tri-plate arrangement and x01= x02= 2 microm and x03= 15 microm lsquorsquo represents the case 2 where the fingers are implemented in a single ended arrangement and x01= x02=x03=2 microm

The tri-plate drive is also implemented in a differential arrangement where the output

is the change in capacitance rather than the total capacitance resulting from an external

excitation This reduces errors cause by off-axis excitation and common mode noise The

output of the differential tri-plate drive is expressed as [39]

Total Top Total BottomC C C∆ = minus where

003 01

1 1( ) (Total TopC N ty

)x x x xε

= +

minus +

(12)

20

002 03

1 1( ) (Total BottomC N ty

)x x x xε

= +

minus +

A graphic comparison of the capacitancedisplacement relation for sensors T1 and T2 is

shown in Figure 3-4 For these gap closing drives it is assumed that for displacements

less that 1 microm the relation is linear indicated by the dashed line above each curve The

sensitivity within this region is approximately 0046 pF microm-1 and 0253 pF microm-1 for T1

and T2 respectively The order of magnitude improvement is attributed to the differential

tri-plate arrangement and thicker comb finger

0

02

04

06

08

1

12

14

16

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-4 lsquorsquo represents the output of the differential tri-plate drive while lsquorsquo represents the output of the transverse drive The dashed line represents the linear approximation for displacements less that 1 microm For the purpose of comparison the output of the transverse drive was equalized to zero at zero displacement

21

In this work the transverse drive is first implemented in sensor T1 using the

PolyMUMPS fabrication process Then in an effort to linearize the sensor response the

lateral drive is used in L1 using SoiMUMPS The reduction in drive sensitivity due to the

lateral drive is further reduced by increasing the gap distance to 3 microm which is

recommended by the foundry to ensure a successful device These reductions are

balanced by increasing the number of comb fingers N and selecting a substrate thickness

that is 125 times large than T1 The third device T2 is fabricated using SoiMUMPS and

uses the tri-plate differential drive This design capitalizes on the increase in sensitivity of

a gap closing drive and larger substrate thickness attaining an order of magnitude

increase in sensitivity over T1 and L1 The specifications for each drive along with the

successive improvement in drive sensitivity are outlined in Table 3-1

Table 3-1 Specifications for each capacitive comb drive implemented in this work T1 L1 and T2

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Comb Finger Thickness tc (microm) 2 25 25 Gap Distance x0 (microm) 2 3 2 Number of Fingers N 200 256 30 Overlap length y0 (microm) 45 10 70 Drive Sensitivity (pF microm-1) 0046 0038 0273

32 MECHANICAL ANALYSIS

321 MICRO-FLEXURE SELECTION

The sensor structure consists of a proof mass connected to a compliant micro-flexure

that itself is anchored to the sensor substrate The proof mass displaces in response to an

22

applied acceleration and this displacement is sensed using the capacitive comb drive

attached to the proof mass

Proof mass

Folded beam micro-flexure

Anchor

Figure 3-5 Sensor Structure

This section focuses on micro-flexure selection as it should exhibit a large linear

displacement range and have a high lateral stiffness ratio to reduce cross-axis sensitivity

A number of micro-flexures have been used in MEMS research these include the

clamped-clamped crab-leg and folded beam shown in Figure 3-6-through 3-7

The clamped-clamped flexure exhibits a high stiffness ratio however for large lateral

displacements extensional axial forces buildup within the beams resulting in a non-linear

force-displacement relation [40] Within the linear regime the stiffness of the clamped-

clamped flexure is expressed as [40]

3

3b

yEt wk

L=

(13)

23

where E is the Youngrsquos Modulus of the beam material (N m-2) tb is the beam out of

plane thickness (m) w is the beam width (m) and L is the beam length (m) The stiffness

ratio is expressed as [40]

2x

y

k Lk w

=

(14)

If L is two orders of magnitude greater than w this results in a ratio of approximately

10000 The linear displacement range is determined using Finite Element Analysis (FEA)

of a clamped-clamped flexure with L=387 microm w=4 microm and tb=2 microm The result of this

simulation is plotted in and compared to a structure with constant stiffness ky

= 035 N m-1 predicted by Eq (13) The linear relation is valid for small deflections that

are less than 2 microm In addition this micro-flexure is susceptible to buckling as a result of

a compressive residual stress that may arise due to the fabrication process [41]

Figure 3-9

In the crab leg design the thigh section is meant to relieve the extensional axial

stress postponing the occurrence of non-linearity until greater lateral displacement [42]

The stiffness of the crab leg flexure is expressed as [43]

3

44

b s s ty

s s t

Et w L LkL L L

κκ

+= +

(15)

where κ=IsIt=(wswt)3 is a ratio of the second moment of inertia of the shin (s) and thigh

(t) ws and wt represent the width and Ls and Lt represent the length of the shin and thigh

regions respectively The stiffness ratio is expressed as [43]

34

4t s s tx

y s t s

w L L Lkk w L L L

αα t

+= +

(16)

The linear displacement range of a crab leg flexure with Ls=285 microm Lt=30 microm ws=wt=3

microm and tb=2 microm is determined using FEA The results are plotted in Figure 3-9 and

compared to a structure of constant stiffness ky=035 N m-1 predicted by Eq (15) The

24

non-linearity is greatly reduced when compared to the clamp-clamped flexure with Eq

accurately predicting the deflection until 4 microm However the flexure is more

susceptible to off-axis deflection as there is an order of magnitude reduction in the

stiffness ratio as compared to the clamped-clamped flexure

(15)

The folded beam flexure stands out as it is designed to exhibit a large linear

displacement range and to relieve residual stress due to the manufacturing process by

allowing the beams to expand and contract along the axial direction [40] It was first used

in an electrostatic resonator and has subsequently been used for electrostatic actuation in

a magnetic hard drives and micro-positioning [3344] If the truss section is rigid

ItrussgtgtIbeam then the stiffness of the flexure is expressed as [4043]

3

32b

yEt wk

L=

(17)

The stiffness ratio is expressed as [40]

2x

y

k Lk w

=

(18)

Figure 3-6 Clamped-clamped flexure

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

20

002 03

1 1( ) (Total BottomC N ty

)x x x xε

= +

minus +

A graphic comparison of the capacitancedisplacement relation for sensors T1 and T2 is

shown in Figure 3-4 For these gap closing drives it is assumed that for displacements

less that 1 microm the relation is linear indicated by the dashed line above each curve The

sensitivity within this region is approximately 0046 pF microm-1 and 0253 pF microm-1 for T1

and T2 respectively The order of magnitude improvement is attributed to the differential

tri-plate arrangement and thicker comb finger

0

02

04

06

08

1

12

14

16

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

Cap

acita

nce

(pF)

Figure 3-4 lsquorsquo represents the output of the differential tri-plate drive while lsquorsquo represents the output of the transverse drive The dashed line represents the linear approximation for displacements less that 1 microm For the purpose of comparison the output of the transverse drive was equalized to zero at zero displacement

21

In this work the transverse drive is first implemented in sensor T1 using the

PolyMUMPS fabrication process Then in an effort to linearize the sensor response the

lateral drive is used in L1 using SoiMUMPS The reduction in drive sensitivity due to the

lateral drive is further reduced by increasing the gap distance to 3 microm which is

recommended by the foundry to ensure a successful device These reductions are

balanced by increasing the number of comb fingers N and selecting a substrate thickness

that is 125 times large than T1 The third device T2 is fabricated using SoiMUMPS and

uses the tri-plate differential drive This design capitalizes on the increase in sensitivity of

a gap closing drive and larger substrate thickness attaining an order of magnitude

increase in sensitivity over T1 and L1 The specifications for each drive along with the

successive improvement in drive sensitivity are outlined in Table 3-1

Table 3-1 Specifications for each capacitive comb drive implemented in this work T1 L1 and T2

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Comb Finger Thickness tc (microm) 2 25 25 Gap Distance x0 (microm) 2 3 2 Number of Fingers N 200 256 30 Overlap length y0 (microm) 45 10 70 Drive Sensitivity (pF microm-1) 0046 0038 0273

32 MECHANICAL ANALYSIS

321 MICRO-FLEXURE SELECTION

The sensor structure consists of a proof mass connected to a compliant micro-flexure

that itself is anchored to the sensor substrate The proof mass displaces in response to an

22

applied acceleration and this displacement is sensed using the capacitive comb drive

attached to the proof mass

Proof mass

Folded beam micro-flexure

Anchor

Figure 3-5 Sensor Structure

This section focuses on micro-flexure selection as it should exhibit a large linear

displacement range and have a high lateral stiffness ratio to reduce cross-axis sensitivity

A number of micro-flexures have been used in MEMS research these include the

clamped-clamped crab-leg and folded beam shown in Figure 3-6-through 3-7

The clamped-clamped flexure exhibits a high stiffness ratio however for large lateral

displacements extensional axial forces buildup within the beams resulting in a non-linear

force-displacement relation [40] Within the linear regime the stiffness of the clamped-

clamped flexure is expressed as [40]

3

3b

yEt wk

L=

(13)

23

where E is the Youngrsquos Modulus of the beam material (N m-2) tb is the beam out of

plane thickness (m) w is the beam width (m) and L is the beam length (m) The stiffness

ratio is expressed as [40]

2x

y

k Lk w

=

(14)

If L is two orders of magnitude greater than w this results in a ratio of approximately

10000 The linear displacement range is determined using Finite Element Analysis (FEA)

of a clamped-clamped flexure with L=387 microm w=4 microm and tb=2 microm The result of this

simulation is plotted in and compared to a structure with constant stiffness ky

= 035 N m-1 predicted by Eq (13) The linear relation is valid for small deflections that

are less than 2 microm In addition this micro-flexure is susceptible to buckling as a result of

a compressive residual stress that may arise due to the fabrication process [41]

Figure 3-9

In the crab leg design the thigh section is meant to relieve the extensional axial

stress postponing the occurrence of non-linearity until greater lateral displacement [42]

The stiffness of the crab leg flexure is expressed as [43]

3

44

b s s ty

s s t

Et w L LkL L L

κκ

+= +

(15)

where κ=IsIt=(wswt)3 is a ratio of the second moment of inertia of the shin (s) and thigh

(t) ws and wt represent the width and Ls and Lt represent the length of the shin and thigh

regions respectively The stiffness ratio is expressed as [43]

34

4t s s tx

y s t s

w L L Lkk w L L L

αα t

+= +

(16)

The linear displacement range of a crab leg flexure with Ls=285 microm Lt=30 microm ws=wt=3

microm and tb=2 microm is determined using FEA The results are plotted in Figure 3-9 and

compared to a structure of constant stiffness ky=035 N m-1 predicted by Eq (15) The

24

non-linearity is greatly reduced when compared to the clamp-clamped flexure with Eq

accurately predicting the deflection until 4 microm However the flexure is more

susceptible to off-axis deflection as there is an order of magnitude reduction in the

stiffness ratio as compared to the clamped-clamped flexure

(15)

The folded beam flexure stands out as it is designed to exhibit a large linear

displacement range and to relieve residual stress due to the manufacturing process by

allowing the beams to expand and contract along the axial direction [40] It was first used

in an electrostatic resonator and has subsequently been used for electrostatic actuation in

a magnetic hard drives and micro-positioning [3344] If the truss section is rigid

ItrussgtgtIbeam then the stiffness of the flexure is expressed as [4043]

3

32b

yEt wk

L=

(17)

The stiffness ratio is expressed as [40]

2x

y

k Lk w

=

(18)

Figure 3-6 Clamped-clamped flexure

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

21

In this work the transverse drive is first implemented in sensor T1 using the

PolyMUMPS fabrication process Then in an effort to linearize the sensor response the

lateral drive is used in L1 using SoiMUMPS The reduction in drive sensitivity due to the

lateral drive is further reduced by increasing the gap distance to 3 microm which is

recommended by the foundry to ensure a successful device These reductions are

balanced by increasing the number of comb fingers N and selecting a substrate thickness

that is 125 times large than T1 The third device T2 is fabricated using SoiMUMPS and

uses the tri-plate differential drive This design capitalizes on the increase in sensitivity of

a gap closing drive and larger substrate thickness attaining an order of magnitude

increase in sensitivity over T1 and L1 The specifications for each drive along with the

successive improvement in drive sensitivity are outlined in Table 3-1

Table 3-1 Specifications for each capacitive comb drive implemented in this work T1 L1 and T2

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Comb Finger Thickness tc (microm) 2 25 25 Gap Distance x0 (microm) 2 3 2 Number of Fingers N 200 256 30 Overlap length y0 (microm) 45 10 70 Drive Sensitivity (pF microm-1) 0046 0038 0273

32 MECHANICAL ANALYSIS

321 MICRO-FLEXURE SELECTION

The sensor structure consists of a proof mass connected to a compliant micro-flexure

that itself is anchored to the sensor substrate The proof mass displaces in response to an

22

applied acceleration and this displacement is sensed using the capacitive comb drive

attached to the proof mass

Proof mass

Folded beam micro-flexure

Anchor

Figure 3-5 Sensor Structure

This section focuses on micro-flexure selection as it should exhibit a large linear

displacement range and have a high lateral stiffness ratio to reduce cross-axis sensitivity

A number of micro-flexures have been used in MEMS research these include the

clamped-clamped crab-leg and folded beam shown in Figure 3-6-through 3-7

The clamped-clamped flexure exhibits a high stiffness ratio however for large lateral

displacements extensional axial forces buildup within the beams resulting in a non-linear

force-displacement relation [40] Within the linear regime the stiffness of the clamped-

clamped flexure is expressed as [40]

3

3b

yEt wk

L=

(13)

23

where E is the Youngrsquos Modulus of the beam material (N m-2) tb is the beam out of

plane thickness (m) w is the beam width (m) and L is the beam length (m) The stiffness

ratio is expressed as [40]

2x

y

k Lk w

=

(14)

If L is two orders of magnitude greater than w this results in a ratio of approximately

10000 The linear displacement range is determined using Finite Element Analysis (FEA)

of a clamped-clamped flexure with L=387 microm w=4 microm and tb=2 microm The result of this

simulation is plotted in and compared to a structure with constant stiffness ky

= 035 N m-1 predicted by Eq (13) The linear relation is valid for small deflections that

are less than 2 microm In addition this micro-flexure is susceptible to buckling as a result of

a compressive residual stress that may arise due to the fabrication process [41]

Figure 3-9

In the crab leg design the thigh section is meant to relieve the extensional axial

stress postponing the occurrence of non-linearity until greater lateral displacement [42]

The stiffness of the crab leg flexure is expressed as [43]

3

44

b s s ty

s s t

Et w L LkL L L

κκ

+= +

(15)

where κ=IsIt=(wswt)3 is a ratio of the second moment of inertia of the shin (s) and thigh

(t) ws and wt represent the width and Ls and Lt represent the length of the shin and thigh

regions respectively The stiffness ratio is expressed as [43]

34

4t s s tx

y s t s

w L L Lkk w L L L

αα t

+= +

(16)

The linear displacement range of a crab leg flexure with Ls=285 microm Lt=30 microm ws=wt=3

microm and tb=2 microm is determined using FEA The results are plotted in Figure 3-9 and

compared to a structure of constant stiffness ky=035 N m-1 predicted by Eq (15) The

24

non-linearity is greatly reduced when compared to the clamp-clamped flexure with Eq

accurately predicting the deflection until 4 microm However the flexure is more

susceptible to off-axis deflection as there is an order of magnitude reduction in the

stiffness ratio as compared to the clamped-clamped flexure

(15)

The folded beam flexure stands out as it is designed to exhibit a large linear

displacement range and to relieve residual stress due to the manufacturing process by

allowing the beams to expand and contract along the axial direction [40] It was first used

in an electrostatic resonator and has subsequently been used for electrostatic actuation in

a magnetic hard drives and micro-positioning [3344] If the truss section is rigid

ItrussgtgtIbeam then the stiffness of the flexure is expressed as [4043]

3

32b

yEt wk

L=

(17)

The stiffness ratio is expressed as [40]

2x

y

k Lk w

=

(18)

Figure 3-6 Clamped-clamped flexure

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

22

applied acceleration and this displacement is sensed using the capacitive comb drive

attached to the proof mass

Proof mass

Folded beam micro-flexure

Anchor

Figure 3-5 Sensor Structure

This section focuses on micro-flexure selection as it should exhibit a large linear

displacement range and have a high lateral stiffness ratio to reduce cross-axis sensitivity

A number of micro-flexures have been used in MEMS research these include the

clamped-clamped crab-leg and folded beam shown in Figure 3-6-through 3-7

The clamped-clamped flexure exhibits a high stiffness ratio however for large lateral

displacements extensional axial forces buildup within the beams resulting in a non-linear

force-displacement relation [40] Within the linear regime the stiffness of the clamped-

clamped flexure is expressed as [40]

3

3b

yEt wk

L=

(13)

23

where E is the Youngrsquos Modulus of the beam material (N m-2) tb is the beam out of

plane thickness (m) w is the beam width (m) and L is the beam length (m) The stiffness

ratio is expressed as [40]

2x

y

k Lk w

=

(14)

If L is two orders of magnitude greater than w this results in a ratio of approximately

10000 The linear displacement range is determined using Finite Element Analysis (FEA)

of a clamped-clamped flexure with L=387 microm w=4 microm and tb=2 microm The result of this

simulation is plotted in and compared to a structure with constant stiffness ky

= 035 N m-1 predicted by Eq (13) The linear relation is valid for small deflections that

are less than 2 microm In addition this micro-flexure is susceptible to buckling as a result of

a compressive residual stress that may arise due to the fabrication process [41]

Figure 3-9

In the crab leg design the thigh section is meant to relieve the extensional axial

stress postponing the occurrence of non-linearity until greater lateral displacement [42]

The stiffness of the crab leg flexure is expressed as [43]

3

44

b s s ty

s s t

Et w L LkL L L

κκ

+= +

(15)

where κ=IsIt=(wswt)3 is a ratio of the second moment of inertia of the shin (s) and thigh

(t) ws and wt represent the width and Ls and Lt represent the length of the shin and thigh

regions respectively The stiffness ratio is expressed as [43]

34

4t s s tx

y s t s

w L L Lkk w L L L

αα t

+= +

(16)

The linear displacement range of a crab leg flexure with Ls=285 microm Lt=30 microm ws=wt=3

microm and tb=2 microm is determined using FEA The results are plotted in Figure 3-9 and

compared to a structure of constant stiffness ky=035 N m-1 predicted by Eq (15) The

24

non-linearity is greatly reduced when compared to the clamp-clamped flexure with Eq

accurately predicting the deflection until 4 microm However the flexure is more

susceptible to off-axis deflection as there is an order of magnitude reduction in the

stiffness ratio as compared to the clamped-clamped flexure

(15)

The folded beam flexure stands out as it is designed to exhibit a large linear

displacement range and to relieve residual stress due to the manufacturing process by

allowing the beams to expand and contract along the axial direction [40] It was first used

in an electrostatic resonator and has subsequently been used for electrostatic actuation in

a magnetic hard drives and micro-positioning [3344] If the truss section is rigid

ItrussgtgtIbeam then the stiffness of the flexure is expressed as [4043]

3

32b

yEt wk

L=

(17)

The stiffness ratio is expressed as [40]

2x

y

k Lk w

=

(18)

Figure 3-6 Clamped-clamped flexure

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

23

where E is the Youngrsquos Modulus of the beam material (N m-2) tb is the beam out of

plane thickness (m) w is the beam width (m) and L is the beam length (m) The stiffness

ratio is expressed as [40]

2x

y

k Lk w

=

(14)

If L is two orders of magnitude greater than w this results in a ratio of approximately

10000 The linear displacement range is determined using Finite Element Analysis (FEA)

of a clamped-clamped flexure with L=387 microm w=4 microm and tb=2 microm The result of this

simulation is plotted in and compared to a structure with constant stiffness ky

= 035 N m-1 predicted by Eq (13) The linear relation is valid for small deflections that

are less than 2 microm In addition this micro-flexure is susceptible to buckling as a result of

a compressive residual stress that may arise due to the fabrication process [41]

Figure 3-9

In the crab leg design the thigh section is meant to relieve the extensional axial

stress postponing the occurrence of non-linearity until greater lateral displacement [42]

The stiffness of the crab leg flexure is expressed as [43]

3

44

b s s ty

s s t

Et w L LkL L L

κκ

+= +

(15)

where κ=IsIt=(wswt)3 is a ratio of the second moment of inertia of the shin (s) and thigh

(t) ws and wt represent the width and Ls and Lt represent the length of the shin and thigh

regions respectively The stiffness ratio is expressed as [43]

34

4t s s tx

y s t s

w L L Lkk w L L L

αα t

+= +

(16)

The linear displacement range of a crab leg flexure with Ls=285 microm Lt=30 microm ws=wt=3

microm and tb=2 microm is determined using FEA The results are plotted in Figure 3-9 and

compared to a structure of constant stiffness ky=035 N m-1 predicted by Eq (15) The

24

non-linearity is greatly reduced when compared to the clamp-clamped flexure with Eq

accurately predicting the deflection until 4 microm However the flexure is more

susceptible to off-axis deflection as there is an order of magnitude reduction in the

stiffness ratio as compared to the clamped-clamped flexure

(15)

The folded beam flexure stands out as it is designed to exhibit a large linear

displacement range and to relieve residual stress due to the manufacturing process by

allowing the beams to expand and contract along the axial direction [40] It was first used

in an electrostatic resonator and has subsequently been used for electrostatic actuation in

a magnetic hard drives and micro-positioning [3344] If the truss section is rigid

ItrussgtgtIbeam then the stiffness of the flexure is expressed as [4043]

3

32b

yEt wk

L=

(17)

The stiffness ratio is expressed as [40]

2x

y

k Lk w

=

(18)

Figure 3-6 Clamped-clamped flexure

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

24

non-linearity is greatly reduced when compared to the clamp-clamped flexure with Eq

accurately predicting the deflection until 4 microm However the flexure is more

susceptible to off-axis deflection as there is an order of magnitude reduction in the

stiffness ratio as compared to the clamped-clamped flexure

(15)

The folded beam flexure stands out as it is designed to exhibit a large linear

displacement range and to relieve residual stress due to the manufacturing process by

allowing the beams to expand and contract along the axial direction [40] It was first used

in an electrostatic resonator and has subsequently been used for electrostatic actuation in

a magnetic hard drives and micro-positioning [3344] If the truss section is rigid

ItrussgtgtIbeam then the stiffness of the flexure is expressed as [4043]

3

32b

yEt wk

L=

(17)

The stiffness ratio is expressed as [40]

2x

y

k Lk w

=

(18)

Figure 3-6 Clamped-clamped flexure

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

25

Figure 3-7 Crab leg flexure

Figure 3-8 Folded beam flexure

The linear range of motion of a folded beam flexure with L=304 microm w=4 microm and tb=2

microm is determined using FEA The results are plotted in Figure 3-9 and compared to a

structure with constant stiffness ky=035 N m-1 predicted by Eq (17) The folded flexure

exhibits a response that closely follows the linear prediction through a displacement that

exceeds the previous two flexures while maintaining the stiffness ratio of the clamped-

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

26

clamped flexure Due to the linear response high stiffness ratio and compensation for

residual stress the folded flexure is implemented in each of the three vibrations sensors

presented in this work

Ideal

Clamped- Clamped

Folded Beam

Crab Leg

00

10

20

30

40

50

60

70

80

00 20 40 60 80 100Applied Force (microN)

Dis

plac

emen

t (microm

)

Figure 3-9 Comparison of the non-linear deflection among the three micro-flexures considered clamped-clamped crab-leg and folded beam All flexures were designed to have the same stiffness of 035 N m-1

322 EQUATIONS OF MOTION

The vibration sensor in this work uses four folded beam flexures in parallel with a

total stiffness kT (N m-1)

3

3

24 bT

Et wk kL

= = (19)

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

27

The sensor can be approximated as a lumped mass system undergoing damped harmonic

oscillation When attached to a vibrating structure the system undergoes base excitation

shown in Figure 3-10 and the equation describing the absolute motion of the mass is

[45]

( ) ( )eff M M B M Bm y c y y k y y+ minus + minusampamp amp amp 0= (20)

where meff is the effective proof mass (kg) yM is the absolute proof mass displacement

(m) yB is the base displacement (m) and c is the coefficient of damping (N s m-1)

Figure 3-10 Lumped mass approximation with base excitation yB(t)

Of specific interest is the relative motion of the mass zy (m) with respect to the base

zy=yM - yB

If a sinusoidal base excitation is applied yB

yB(t)=YB sin(ωt)

where YB is the base excitation amplitude (m) ω is the base excitation frequency (rad s-1)

and t represents time the equation governing the relative motion of the proof mass is

expressed as [45]

2 22 sy n y n y Bz z z Y in( )tζω ω ω ω+ + =ampamp amp (21)

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

28

where ζ=05b(km)12 is the damping ratio ωn=(km)12 is the proof massflexure natural

frequency (rad s-1) and ω is the base excitation frequency (rad s-1) Eq (21) shows that

the system can be treated as a case of direct mass excitation shown in Figure 3-11 where

the acceleration applied to the mass a (m s-2) is

a=ω2YB sin(ωt)

Figure 3-11 Lumped mass approximation with direct mass excitation z(t)

By evaluating the Laplace transform of the left side of Eq (21) and making the

substitution s=jω the magnitude of the system transfer function is expressed as

2 2 2

1( )( ) (2n n

H jω2)ω ω ζω ω

=minus +

(22)

In the low frequency range ωltltωn the system response is well approximated from the

DC response

H(ω=0)=1ωn2

For a low frequency external acceleration of peak amplitude A (g) the peak response Zy

(m) is expressed as

2

981y

n

AZωsdot

= (23)

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

29

Therefore by decreasing the system natural frequency the response to a low frequency

excitation can be increased The decrease in ωn can be accomplished by decreasing the

flexure stiffness or increasing the proof mass The stiffness can be decreased by

increasing the flexure length however depending on the fabrication process a maximum

length is placed on beams due to out of plane bending Increasing the proof mass can be

accomplished by increasing the proof mass area However limitations are placed on this

approach which are dependent on the fabrication process For example for overall device

mechanical stability in SoiMUMPS a restriction is placed on the amount of removed

bulk material which is required to free moving structures This places a restriction on the

maximum area of the proof mass

A FEA of the three proposed sensors are done to accurately determine the natural

frequency and flexure stiffness of the structures Table 3-2 outlines the sensor

dimensions and mechanical properties The stiffness and natural frequency were found to

be 14 N m-1 and 3500 Hz 378 N m-1 and 1240 Hz and 433 N m-1and 750 Hz for T1

L1 and T2 respectively and represent a good match with the analytical formula

The effectiveness of the proof massflexure mechanical system can be quantified by

considering the mechanical sensitivity SM (m g-1) is given by

2

981yM

n

VS

A ωpart

= =part

(24)

SM is evaluated to be 002 microm g-1 0161 microm g-1 and 0442 microm g-1 for T1 L1 and T2

respectively Sensors L1 and T2 attain an order of magnitude increase over T1 which is

attributed to the thicker substrate used in SoiMUMPS The nearly three fold increase

from L1 to T2 is a result of a substantially larger proof mass area in T2

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

30

T1 L1 T2 Fabrication Process PolyMUMPS SoiMUMPS SoiMUMPS Beam Length L (microm) 650 1048 1000 Beam Width w (microm) 4 8 8 Beam Thickness tB (microm) 2 25 25 Youngrsquos Modulus E (GPa) 158 169 169 Natural Frequency ωn (Hz) 3500 1240 750 Flexure Stiffness k (N m-1) 14 378 433 Effective Mass meff (kg) 29e-9 56e-8 195e-7 Mechanical Sensitivity (microm g-1) 002 0161 0442

Table 3-2 Dimensions and mechanical properties for the three MEMS vibration sensors outlined in this work

323 QUALITY FACTOR

The quality factor Q is the ratio of total energy stored in the system to the energy

lost per cycle The total quality factor Qtotal is expressed as [3]

s

1 1 1

total f mQ Q Q Q= + +

1 (25)

where 1Qf represents the energy lost to the surrounding fluid 1Qs is the energy lost

through the supports and 1Qm is the energy lost through the sensor material For MEMS

sensors Qf is the dominant component and is the focus of the discussion A prediction of

the quality factor is important as it influences the mechanical noise of the sensor

For MEMS devices operated in ambient air energy is lost to the surrounding fluid

through air damping The quality factor as a function of air damping is expressed as

T

n total

kQcω

= (26)

where ctotal is the total damping coefficient (N s m-1) The effects of air damping are

approximated by two mechanism slide film and squeeze film damping Slide film

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

31

damping is caused by the fluid in the gap between two parallel moving plates If a linear

velocity profile is assumed the coefficient of damping cslide-film is approximated by

air slide filmslide film

Acd

micro minusminus =

(27)

where microair=18e-5 N s m-2 is the absolute viscosity of air at 20degC Aslide-film is the effective

area between the two plates (m) and d is the distance between the two plates (m)

Squeeze film damping arises when the gap between two parallel plates decrease

squeezing the fluid between the two plates creating an opposing force Squeeze film

damping csqueeze-film can be approximated using the following relation [3846]

3

3air plate plate

squeeze film

L Wc

dmicro α

minus = (28)

where Lplate is the effective plate length Wplate is the effective plate width and α is a

parameter dependant on the WplateLplate ratio which can be determined from reference

[46] For the sensor T1 slide film damping is assumed between the proof mass and

substrate while squeeze film damping is assumed between the fixed and movable fingers

on the transverse comb drive The total damping coefficient ctotal is expressed as the sum

of the individual contributions

total slide film squeeze filmc c cminus minus= + (29)

For sensor L1 only slide film damping is assumed between the fixed and movable

fingers on the lateral comb drive For sensor T2 only squeeze film damping is assumed

between the fixed and movable fingers on the tri-plate differential comb drive

summarizes the damping constant and approximate quality factors for each sensor

Table 3-3

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

32

Sensor ctotal (N s m-1) Q T1 617e-7 103 L1 468e-7 631 T2 171e-5 53

Table 3-3 Summary of the theoretical coefficients of damping and quality factors for the three sensors presented in this work

33 DRIVE STABILITY

Gap closing capacitive drives such as T1 and T2 suffer from pull-in instability that

places a limit on drive displacement Pull-in occurs since capacitive readout electronics

apply a voltage driving signal across the fixed and moving comb drive which creates an

electrostatic force that brings the fingers together On a lateral drive this force is

balanced at rest and after a displacement however in gap closing drives the force is only

balanced at rest The voltage signal is time varying with peak VE and is typically a square

wave with a 50 percent duty ratio or sinusoidal and has a frequency a few orders of

magnitude higher than the proof massflexure natural frequency [47] For sensors T1 and

T2 the total electrostatic force Fx-Elect applied to the proof mass due to the capacitive

drives is expressed as

2 1

12

Drivex Elect

CF Vx

part=

part

(30)

where V1=VEradic2 for a sinusoidal voltage signal x is the displacement direction

CDrive=CTotalT for T1 and CDrive=CTotalTop+ CTotalBottom for T2 The electrostatic force

applied on T1 and T2 is expressed

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

33

20

1 2 20 0

1 12 ( ) ( )x Elect T

N ty VFx x x x

εminus

= minus minus +

20

2 2 2 203 01 02 03

1 1 1 12 ( ) ( ) ( ) ( )x Elect T

N ty VFx x x x x x x x

εminus

= minus + minus

minus + minus + 2

(31)

Neglecting the effect of damping the equation governing the motion of the proof mass is

expressed as

eff x Electm a kx F= minus (32)

To determine the pull-in acceleration and displacement Eq (32) was solved in [47] by

plotting the right hand side (RHS) for various x values and then plotting the left hand

side (LHS) as a horizontal line for various applied accelerations The RHS of Eq (32) is

plotted in for T1 and T2 and it is observed that there are two points of

intersection between any horizontal line and the RHS curve The first point of

intersection represents the stable solution while the second the unstable The maximum

point of the RHS curve represents the maximum acceleration that can be applied for a

stable response This is approximately 42g for R1 and 125g for T2 beyond these values

the undesired pull-in effect occurs

Figure 3-12

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

34

000

050

100

150

200

250

300

350

400

450

0 02 04 06 08 1 12 14 16 18

Drive Displacement (microm)

RH

S (micro

N)

Figure 3-12 Graph representing stability regions for gap closing capacitive drives lsquorsquo represents the RHS of Eq (32) for T2 while lsquorsquo represents the RHS of T1

34 ELECTROSTATIC ACTUATION

The sensor T1 is fabricated with an electrostatic lateral comb drive which is used to

actuate the proof mass and allow for verification of the sensor frequency response The

electrostatic force generated when a voltage is applied across a parallel plate capacitor is

expressed as [28]

212

Total Lact act

CF V

ypart

=part

(33)

where Fact is the actuation force in the y-direction CTotalL is the capacitance of the lateral

comb drive expressed by Eq (10) and Vact is the applied actuation voltage across the

capacitor plates (v) For the lateral comb drive shown in Figure 3-1 the force in the y

direction is expressed as

2

0act act

N tF Vxε

= (34)

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

35

A sinusoidal excitation voltage with frequency ωv results in an applied force with a DC

component and a time varying components with frequency 2ωv

35 SECTION SUMMARY

Table 3-4

Table 3-4 Summary of the characteristics of the sensors outlined in this work

summarizes the important characteristics of the three sensors presented in

this work Changes are made in each sensor that progressively improves the mechanical

and capacitive drive sensitivity Since T1 and T2 use gap closing capacitive drives they

are subjected to excitation limitations to maintain stability In addition T2 offers a more

linearized response due to the tri-plate arrangement

T1 L1 T2 Mechanical Sensitivity (microm g-1) 002 0161 0442 Capacitive Drive Sensitivity (pF microm-1) 0046 0038 0253 Maximum Excitation (g) 42 na 12g Maximum Excitation for Linearized Response (g)

39 na 115

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

4 CAPACITANCE TO VOLTAGE READOUT CIRCUIT

Capacitance measurements can be made by applying an AC excitation signal on one

node of a capacitor and detecting the flow of charge using a charge amplifier as shown

in Figure 4-1 The amplitude of the output voltage VCA (V) is determined by the feedback

capacitor Cf and the amplitude of the excitation signal VE and is expressed as [48]

sensorCA E

f

CV VC

= sdot (35)

Figure 4-1 Ideal charge amplifier

This basic charge amplifier has been used as a building block for capacitance to

voltage readout circuitry for many capacitance based sensors In an off-chip

implementation it has been used as a readout for an open-loop accelerometer in [49] and

for a MEMS based power disconnect in [38] In the circuit utilized in this work shown in

R1 is required to keep the proof mass of the MEMS sensor at a defined

potential R2 prevents the output from drifting and saturating the amplifier and C1 blocks

any DC component in the sinusoid at the negative input of op-amp 1 [49] If R1 Rf and C1

are selected to be large ie 5 MΩ 2 MΩ and 1 nF respectively the output of the

charge amplifier can be closely approximated by Eq (35) In this work the impedances

Figure 4-2

36

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

37

at the negative input of op-amp 1 were balanced at the positive input by Cb and Rb to

reduce the voltage offset caused by the amplifierrsquos input bias current and to improve the

noise performance [5051]

Figure 4-2 Capacitance to voltage readout circuit which is made up of four

elements the charge amplifier voltage amplifier demodulator and low pass filter

Overall the circuit consists of four elements the charge amplifier voltage amplifier

demodulator and a RC low pass filter The output of the charge amplifier is followed by

a variable gain voltage amplifier with output VVA

VVA= VCAR3R2

This voltage is then directed to the demodulator with output VDM expressed as

VDM=VVAVE10

The output of the demodulator is then low pass filtered with cut-off frequency ωc of

1591 Hz to yield an approximate DC signal The output of the demodulator is a positive

AC signal with a DC offset with the low pass filter effectively removing the AC

component The output of the low pass filter VLP is approximately

VLP = VDM 2

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

38

The DC signal VLP represents the rest capacitance of the sensor as determined by the

comb drive

The circuit was simulated using Orcad and the signals at various nodes are shown in

and The output of the readout circuitry at the low pass filter is

closely approximated by

Figure 4-3

Figure 4-3 Orcad simulation results of the capacitance to voltage readout circuit

The excitation voltage VE is 75 VPK and the output of the voltage amplifier VVA is

approximately 13125 VPK

Figure 4-4

23

220E

LP sensorf

V RV CR C

= sdot (36)

The sensitivity of the readout circuit SRO (V pF-1) is expressed as

23

220ELP

ROsensor f

V RVSC Rpart

= =part C

(37)

With the components shown in Figure 4-2 SRO is approximately 141 V pF-1

15V VVAVE 10V

0

-10V

Time

400us -15V

430us 435us 405us 410us 415us 425us 440us415us

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

39

Figure 4-4 Orcad simulation results showing the sinusoidal output of the

demodulator VDM with a 984 VPK-PK amplitude and the DC output of the low pass

filter VLP at approximately 5 V

Time

400us 410us 415us 425us 430us 435us 440us

0

-15V

VDM 15V 10V VLP

-10V

405us 420us

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

5 NOISE ANALYSIS

51 MECHANICAL-THERMAL NOISE

For small signal sensors the micro-structures are susceptible to mechanical noise

arising from collisions of the surrounding fluid molecules [52] This noise applies a

fluctuating force on the proof mass with spectral density [52]

4N BF k= Tc (38)

where FN is the noise force (NradicHz) kB=138times10-23 JK is Boltzmannrsquos constant and T is

the absolute temperature (K) The mechanical noise displacement is then expressed as

[4552]

22 2

4

1

B totalN

Tn n

k TcX

f fkf f

= minus + Q

(39)

where XN is the noise displacement (mradicHz) kT is the flexure spring constant f is the

excitation frequency (Hz) and fn is the system natural frequency (Hz) If f is much less

than fn then using Eq (23) the noise equivalent acceleration an (gradicHz) is well

approximated by

41981

B nn

k TamQ

ω=

(40)

Therefore to reduce the mechanical noise the sensor should have a large proof mass This

can be achieved using a large area and a fabrication process with a thick structural layer

A high quality factor is also required however MEMS sensors operated in air suffer from

high damping which greatly reduces Q which could be overcome by vacuum packaging

40

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

41

The noise equivalent accelerations for the sensors outlined in this work are 35 microg Hz-frac12

019 microg Hz-frac12 and 027 microg Hz-frac12 for T1 L1 and T2 respectively The order of magnitude

decrease in L1 and T2 over T1 is due to the increase in proof mass size and the

corresponding decrease in the natural frequency

52 ELECTRICAL NOISE

In active devices the primary sources of noise are voltage and current noise which

appear as flicker noise (1f) at low frequency and white noise high frequency while

thermal noise exists in the resistors as white noise [51] To measure the small variations

in capacitance of the MEMS device the noise sources are included in the analysis A

noise analysis is done in PSpice using the AD712 operational amplifier model provided

by Analog Devices The model adds a voltage source in series with the positive input to

include the effect of the input noise voltage and a current source between each input

(positive and negative) and ground to include the effect of the input noise currents As

well Orcad includes the thermal noise of resistors in the analysis At the node VVA of the

circuit shown in Figure 4-2 the noise voltage spectral density eno is shown in

The total output noise voltage (RMS) at the output is the square root of the area

under the noise voltage power spectrum eno2 and is shown in Figure 5-2 Since the

circuit has a low pass filter with a cutoff frequency of 1591 Hz the total voltage noise

Eno is approximately 30 microVrms

Figure

5-1

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

42

Figure 5-1 Orcad simulation of the noise voltage spectral density eno at the output of the voltage amplifier

Figure 5-2 Total output noise Eno at the voltage amplifier which is expressed as Eno=( int eno

2)12

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

6 FABRICATION OVERVIEW

This section briefly overviews the two foundry fabrication processes that are used for

the sensors developed in this work [5354] The first is PolyMUMPS which is a three

layer polysilicon surface micromachining process The second is SoiMUMPS which is a

bulk micromachining process that uses silicon as the structural layer

61 POLYMUMPS

In this work only the first two polysilicon layers POLY0 and POLY1 are used in the

T1 sensor layout Fabrication of the MEMS sensor begins on top of a silicon wafer that is

doped with phosphorus to prevent charge feed-through and coated with a layer of silicon

nitride for electrical isolation shown in Figure 6-1 A The first polysilicon layer POLY0

is deposited at a thickness of 05 microm and is subsequently patterned using lithography

and etched using reactive ion etching (RIE) as shown in Figure 6-1 B POLY0 is a non-

mechanical layer and in this work is used as a ground plane Next a phosphosilicate

glass (PSG) layer which acts as a sacrificial layer is deposited using low pressure

chemical vapor deposition (LPCVD) then patterned using lithography and etched RIE

as shown in Figure 6-1 C The PSG allows for the creation of suspended structures and

its thickness of 2 microm sets the distance between the suspended structures and the ground

plane In areas where the PSG is etched away the suspended structures are anchored

Next the first structural polysilicon layer POLY1 is deposited at a thickness of 2 microm

patterned and etched as shown in Figure 6-1 D POLY1 represents the mechanical layout

of the sensor as it defines the flexures proof mass and the fixed and movable comb

fingers

43

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

44

Figure 6-1 PolyMUMPS fabrication process for a simple cantilever beam

62 SOIMUMPS

The MEMS sensor is fabricated using a silicon-on-insulator (SOI) wafer that consists

of four layers the silicon structural layer insulating oxide substrate and bottom-side

oxide as shown in Figure 6-2 A To define the sensor structures the silicon layer is

patterned using lithography and etched using deep reactive ion etching (DRIE) as shown

in Figure 6-2 B Next the bottom side oxide substrate and insulating oxide are patterned

using lithography and etched using RIE DRIE and wet etching respectively as shown

in Figure 6-2 C This step is required over any area where the structural layer is required

to be movable In this work the substrate needs to be removed under the flexures proof

mass and movable comb fingers The silicon structural layer is available in thicknesses

of 10 microm or 25 microm

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

45

Figure 6-2 SoiMUMPS fabrication process for a simple cantilever beam

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

7 EXPERIMENTAL RESULTS

71 MEMS SENSOR ndash T1

The sensor T1 is fabricated first to explore device sensitivity response to excitation

and possible design errors The 45times45mm sensor die is attached and wire bonded to a

68PGA socket to facilitate testing First the sensor is mounted on a six degree of freedom

robotic manipulator with an attached 01 microm probe tip to evaluate the capacitance-

displacement relation Shielded coaxial cables are soldered to the two pins of the variable

capacitor and connected to a capacitance to digital readout board (Analog Devices AD

7746) Then using the probe tip the proof mass is moved in increments of approximately

02 microm Figure 7-1 shows the experimental data points plotted along with the theoretical

capacitancedisplacement curve according to Eq (9) The experimental data points follow

the theoretical curve well however for displacements greater that 12 microm instability

resulting in greater deviation from the theoretical prediction is observed This is

attributed to the pull-in effect which limits the displacement range of the transverse

comb drive For displacements less than 1 microm an average drive sensitivity of 0024 pF

microm-1 is observed and is indicated by a dashed line and the overall device sensitivity is

000048 pf g-1 This represents a decrease of 54 from the theoretical value of 000092

pF g-1 This discrepancy is attributed to parasitic capacitance which affects the sensor

response in two ways it adds a DC offset and it minimizes the capacitive drive

sensitivity The parasitic capacitance arises between the sensor and substrate the sensing

lines on the 68PGA package bonding pads and substrate and any wiring in the readout

circuit For example large bonding pads on the order 100micromtimes100microm (the size used in

46

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

47

T1) which are required to simplify the wire bonding process can add about one pico-

farad of capacitance to the sensor [55] This degrades the sensitivity of the device which

is in the femto-farad range Parasitic capacitance can be greatly reduced by packaging the

sensor and readout together eliminating the need of long coaxial cables and other wiring

or ideally by fabricating both on the same substrate which eliminates the requirement of

bonding pads

0

01

02

03

04

05

06

07

08

0 02 04 06 08 1 12 14 16 18 2

Displacement (microm)

Capa

cita

nce

(pF)

Figure 7-1 The solid line represents the theoretical capacitance versus displacement relation while the solid points represent the experimental values The dashed line represents the assumed linear relation for displacements less than 1microm The experimental average drive sensitivity is approximately 0024 pF microm-1 For displacements greater than 12 microm drive instability is observed and is attributed to the pull-in effect

Next to evaluate the mechanical sensitivity the electrostatic comb drive is used to excite

the flexureproof mass system into resonance to determine the sensor natural frequency

The sensor was mounted inside a custom fabricated vacuum chamber to reduce the effect

of air damping on the response [56] The frequency of the sinusoidal voltage signal

excitation signal is varied in increments of 200 Hz and the amplitude and phase of the

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

48

output signal is shown in Figure 7-2 The resonance peak is located at approximately

3600 Hz and is accompanied with a phase shift of 168deg The experimental mechanical

sensitivity as determined by Eq (24) is 0019 microm g-1

0

05

1

15

2

25

3

35

4

45

400 800 1200160020002400280032003600400044004800520056006000

Frequency (Hz)

Am

plitu

de (V

)

-80

-60

-40

-20

0

20

40

60

80

100

120

Phas

e (deg

)

Figure 7-2 lsquotimesrsquo represents the amplitude and solid line represents the phase response The resonance peak is located at 3600 Hz and accompanied by a phase shift of 168deg

Next the sensor is mounted near the free end of a clamped-free aluminum beam with

a fundamental natural frequency of approximately 23 Hz By displacing the beam free

end and releasing it accelerations in the range of 0-20 g are generated The accelerations

are verified using a commercial accelerometer Model 352C67 manufactured by PCB

Piezotronics At rest the sensor produces a RMS capacitance of 0014 pF which

represents the noise in the system When divided by the experimental device sensitivity

of 000046 pf g-1 this corresponds to a noise equivalent acceleration of 30g Therefore

for applied accelerations less than 30g the sensor produces no observable response as the

minute displacements are buried in the noise signal For an applied acceleration of 20 g

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

49

the theoretical prediction estimated a displacement of 040 microm resulting in a capacitance

increase of 10 fF The lack of a distinguishable response is due to the low capacitive

drive and mechanical sensitivity which both can be attributed to the minimal 2microm

structural layer thickness The mechanical sensitivity could be increased by enlarging the

proof mass area however since the out of plane stiffness is of the same order of

magnitude as the primary direction this may lead to shorting between the structural layer

and the ground plane which are only 075 microm apart

The limitations of the sensor are found to be dependant on the fabrication process

and the integration of the readout circuit and sensor The primary fabrication limitation is

the 2 microm structural layer thickness of PolyMUMPs which impedes the capacitive drive

and mechanical sensitivities In addition the excessive wiring required to connect the

AD7746 readout board to the sensor contributes to a large parasitic capacitance that

further reduces the device sensitivity The fabrication limitation can be reduced by

selecting the SoiMUMPS process which provides a structural layer thickness of 25 microm

In addition the parasitic capacitance can be reduced by using a custom capacitance to

readout circuit mounted close to the 68PGA socket eliminating the AD7746 board and

long wiring

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

50

Figure 7-3 A SEM micro-graph showing the sensor TI fabricated using PolyMUMPS The four folded beam flexures had a length of 650 microm and total stiffness of 14 Nm The effective proof mass was 29e-9 kg and the system had a natural frequency of 3500 Hz

Figure 7-4 A SEM micro-graph of the transverse comb drive implemented in T1 with 200 comb fingers a gap distance of 2 microm overlap length of 45 microm and thickness of 2 microm The average experimental sensitivity was found to be 0024 pF microm-1 for displacements of less than 1 microm

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

51

72 READOUT CIRCUITRY

The readout circuitry is fabricated using discrete components on a prototype board

The charge amplifier and voltage amplifier are mounted as close as possible to the

68PGA socket to minimize electronic noise and parasitic capacitance Due to space

constraints the multiplier and low pass filter are placed on a separate prototype board In

addition to the circuit components of 01 microF and 1 microF capacitors are placed in

parallel to the power supplies of each IC component to minimize any high frequency

noise

Figure 4-2

An input excitation signal VE of 75 VPK at a frequency of 100 kHz is applied to the

circuit and all IC components were powered with plusmn15 VDC supply voltages The readout

circuit is first calibrated using discrete capacitors from 1-4 pF and the sensitivity is found

to be 13 V pF-1 compared to the theoretical sensitivity of 141 V pF-1 The calibration

curve is shown in Figure 7-5 The discrepancy in the output voltage is attributed to the

capacitors which could differ by plusmn025 pF from their ideal values The MEMS sensor is

then connected to the readout circuit and an output voltage of 49 V was observed which

corresponds to a total capacitance of 35 pF

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

52

0

1

2

3

4

5

6

0 05 1 15 2 25 3 35 4

Capacitance (pF)

Out

put V

olta

ge (V

)

Figure 7-5 The solid line represents the theoretical capacitance to voltage relation expressed by Eq (37) with a sensitivity of 141 V pF-1 The solid points represent the experimental data points when discrete capacitors 1-4 pF are used in the readout circuit The horizontal dashed line represents the output voltage of the MEMS sensor at rest and corresponds to a capacitance of 35 pF

73 MEMS SENSORndash L1

The sensor L1 is fabricated using the SoiMUMPS process and tested with the analog

capacitance to voltage readout circuit First with the MEMS vibration sensor at rest the

output of the readout circuitry at various stages is measured using an Agilent 54600

oscilloscope The top trace in represents the input excitation signal from a BK-

Precision 4011a Function Generator while the bottom trace represents the output of the

voltage amplifier The output of the demodulator and the low pass RC filter is shown in

The results exhibit a slight variation from the theoretical values which is

attributed to non-ideal capacitors and resistors

Figure 7-6

Figure 7-7

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

53

Figure 7-6 Oscilloscope trace of the input excitation voltage VE asymp 75 VPK and output of the voltage amplifier VVA asymp 103 VPK Readings were taken with the MEMS sensor at rest

Figure 7-7 Oscilloscope trace of the demodulator output VDM asymp 76 VPK-PK and output of the low pass filter VLP asymp 36 V Readings were taken with the MEMS sensor at rest

Next the sensor and readout circuit are mounted on the six degree of freedom robotic

manipulator to determine the capacitive drive sensitivity The proof mass is displaced in

increments of 02 microm and the output voltage of the readout circuitry at the low pass filter

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

54

VLP for each step is recorded Figure 7-8 displays the capacitance values referenced to

zero which are determined using the experimental readout circuit sensitivity value

0

005

01

015

02

025

03

035

04

0 1 2 3 4 5 6 7 8 9 10Displacement (microm)

Cap

acita

nce

(pF)

Figure 7-8 lsquorsquo represents the experimental data points with the dashed line representing the fitted curve Solid line represents the theoretical data points according to Eq (10)

The experimental sensitivity for the lateral drive is found to be 0025 pF microm-1 and

represents a 32 decrease from the theoretical value of 0038 pF microm-1 The discrepancy

is attributed to parasitic capacitance that is inherent to the sensor and packaging

Next the dynamic response of the sensor is tested by mounting it near the free end of

a clamped-free aluminum beam The output of the readout circuitry at the low pass filter

is connected to an Agilent 35670 Dynamic Signal Analyzer The experimental setup is

shown in Figure 7-9 and Figure 7-10 To capture only the AC change in the output

voltage the signal analyzer is set to AC coupling to remove the DC offset

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

55

Function Generator Power

Supplies

Readout (DM amp LP)

Oscilloscope

Dynamic Signal Analyzer Clamped-

Free Beam MEMS Sensor and Readout (CA amp VA)

Figure 7-9 Experiment setup which includes the MEMS senor and readout board mounted on the clamped-free beam and test equipment

MEMS Sensor

68 PGA Package

Commercial Accelerometer

Figure 7-10 Close-up of the MEMS sensor and readout board mounted on the test beam along with the commercial piezoelectric accelerometer Only the charge amplifier (CA) and the voltage amplifier (VA) are mounted on the same prototype board as the MEMS sensor due to space constraints

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

56

To calibrate the MEMS sensor the results are compared with the true beam

acceleration which is determined by the commercial piezoelectric accelerometer

mounted near the beam free end as shown in Figure 7-10 This accelerometer is factory

calibrated and exhibits a sensitivity of 01 V g-1 and resolution of 016 mg As the beam

is set into vibration the response of each sensor is shown in Figure 7-11 The dynamic

response of the MEMS sensor matches the piezoelectric accelerometer exhibiting the

expected exponentially decreasing sinusoidal beam vibration

Figure 7-11 Figure comparing the response of a commercial piezoelectric accelerometer and MEMS sensor (L1) to a beam vibration A) Dynamic response of the piezoelectric accelerometer representing the true beam vibration with a peak voltage 114 V corresponding to an applied acceleration of 114 g B) Dynamic response of the MEMS sensor with a peak voltage of 006 V

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

57

Next as the beam is set into vibration two traces are exported from the signal

analyzer and are shown in The top trace represents the time domain signal of

the beam vibration while the bottom represents the frequency spectrum The maximum

voltage VLPMAX of the time domain signal is 73 mV which corresponds to acceleration

of approximately 17 g The frequency spectrum shows a peak at approximately 23Hz

which corresponds to the fundamental natural frequency of the beam

Figure 7-12

Figure 7-12 Sensor dynamic response to the vibration of the clamped-free beam and represents a direct export from the dynamic signal analyzer (Agilent 35670) Top trace is the time domain signal of the low pass filter output where VLPMAXasymp73mV which represents a maximum acceleration of approximately 15g The bottom trace represents the frequency spectrum with a peak at 23Hz which corresponds to the beam natural frequency

The maximum output voltage VLPMAX is measured for six accelerations tests

conducted within a range of 585-18 g The linear response of the sensor and readout

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

58

circuitry is shown graphically in Figure 7-13 The theoretical sensitivity S (mV g-1) is

found to be 862 mV g-1 and is expressed as

RO L MS S S S= times times (41)

The experimental sensitivity for the sensor which represents the slope of the linear curve

fitted to the experimental data is 4258 mV g-1 and represents a 50 difference from the

theoretical value This is largely attributed to the presence of parasitic capacitance which

can be further reduced by incorporating the MEMS device and circuit components on the

same substrate using a CMOS fabrication process The decrease in sensitivity is also

attributed to manufacturing irregularities that result in variations to the ideal mechanical

characteristics such as beam dimensions and Youngrsquos Modulus The discernable signal

from L1 can be attributed to its mechanical sensitivity which represents an order of

magnitude increase over T1

0

001

002

003

004

005

006

007

008

009

01

0 5 10 15 20

Input Acce leration (g)

Out

put V

olta

ge (V

)

Figure 7-13 Experimental results showing the response of the sensor to various accelerations in the range of 0-20 g where lsquorsquo represents the experimental data points and the solid line represents the fitted curve The experimental sensitivity was found to be 4258 mV g-1

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

59

The minimum detectable acceleration (g) is determined by the Brownian noise of the

MEMS sensor and the electronic noise of the readout circuitry along with the device

sensitivity and is found by analyzing the time domain signal with the sensor at rest

Min Det Acceleration = RMS Noise Voltage Device Sensitivity

This resulted in a RMS voltage of 1658 mV and represents an equivalent acceleration of

038g This value sets the detection limit for the sensor and only accelerations above this

value can be detected The noise voltage is two orders of magnitude higher than the

predicted value of 003 mV and the increase is attributed to 60 Hz electronic noise from

the surrounding test environment This can be reduced by completely shielding the

MEMS sensor and readout circuitry and noise from the function generator and power

supplies By comparison in a bandwidth of 1591 Hz a CMOS-MEMS sensor developed

in [57] exhibited a minimum detectable acceleration of 004 g an order of magnitude

lower than L1 In the same bandwidth an Analog Devices accelerometer ADXL103

fabricated using their proprietary Integrated Microelectromechanical Systems (iMEMS)

process that combines the sensor and signal conditioning circuitry achieves a minimum

detectable acceleration of 00044 g two orders of magnitude lower that L1

74 SECTION SUMMARY

This section overviewed the experimental results for two MEMS sensors that are

fabricated and tested The first sensor T1 did not achieve a distinguishable response

which is attributed to the low mechanical sensitivity and the influence of parasitic

capacitance brought on by the MEMS device packaging and wiring The second sensor

L1 is fabricated in the SoiMUMPS process and achieved good results as the thicker

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

60

structural layer improved the capacitive drive and mechanical sensitivities in turn

increasing the overall device sensitivity summarizes the experimental results

for each sensor

Table 7-1

Table 7-1 Summary of experimental results

T1 L1 Fabrication PolyMUMPS SoiMUMPS Capacitive Drive Sensitivity (pF microm-1) 0024 0025 Device Sensitivity (g mV-1) na 4258 Noise Floor (g Hz-frac12) na 00098 Min Detectable Acceleration (g) 30 038

The sensor noise floor which determines the minimum detectable acceleration is

determined by the Brownian noise of the MEMS sensor and the electronic noise of the

readout circuit The latter is the dominant component and represents the limit for off-

chip readout circuitry For the noise floor to reach the physical limit of Brownian noise

the MEMS sensor and readout circuitry must be incorporated on the same substrate using

an integrated MEMS-CMOS process and be well shielded

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

8 CONCLUSION AND FUTURE WORK

81 CONCLUSION

This work presented three MEMS capacitive sensors for the detection of low

frequency vibrations The first sensor T1 used the PolyMUMPS foundry fabrication

process while the latter two L1 and T2 used SoiMUMPS Experimental results are

presented for T1 and L1 and the experiments conducted include static tests on a six

degree of freedom robotic manipulator with a 01 microm probe tip and dynamic tests on a

vibrating cantilever beam Capacitance measurements are initially made using a

commercial capacitance to digital readout board (AD 7746) In an attempt to reduce

parasitic capacitance and environmental electronic noise due to the wiring required to

connect the MEMS package to the AD 7746 an analog capacitance to digital readout

board is fabricated using discrete components

Experimental results for T1 produced a transverse comb drive sensitivity of 0024 pF

microm-1 that is 54 less than the predicted value The discrepancy could be due to fringe

capacitances that are neglected in the theoretical calculations parasitic capacitance that

arise between the sensor and substrate sensing lines on the packaging and wiring and

any out of plane deflection cause by residual stress that results in a misalignment between

the comb drive fingers In addition the sensor produced no observable response when

subjected to vibration in the range of 0-20g as the minute displacements are buried in

noise This is due to the low sensitivity of the device which is limited by the substrate

thickness of 2 microm in PolyMUMPS Increasing the thickness effects the achievable natural

frequency and comb drive capacitance and results in an increase in overall device

61

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

62

sensitivity This is addressed by selecting the SoiMUMPS process which has a structural

layer thickness of 25 microm and the potential to increase device sensitivity by an order of

magnitude The sensor L1 is also designed with a lateral comb drive that has an intrinsic

linear response In addition the fabricated sensor is tested with a custom made

capacitance to voltage readout circuit The experiments resulted in a capacitance drive

sensitivity of 0025 pF microm-1 that is 32 lower than the predicted value which implies

that parasitic capacitance is still an important issue The sensor exhibited a sensitivity of

4258 mV g-1 when subjected to cantilever beam vibrations between 0-20g and a

minimum detectable acceleration of 038g

The third MEMS sensors T2 is designed and submitted for fabrication using the

SoiMUMPS process It increases sensitivity by two orders of magnitude over L1 to an

estimated 170 mV g-1 using a tri-plate drive and is implemented in a differential

arrangement to reduce noise However ultimately the device performance is limited by

parasitic capacitance that reduces device sensitivity and electronic noise that increase the

minimum detectable acceleration

82 FUTURE WORK

Future work should investigate the use of a foundry CMOS process to implement the

MEMS sensor and readout electronics This can be done in two ways

bull MEMS sensor and readout circuit are fabrication using two different

processes For example SoiMUMPs for MEMS and a 08 micron CMOS

process for the readout which are then combined using wafer level

packaging

bull MEMS sensor and readout circuit are fabricated on the same substrate

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

63

In addition future work should also focus on developing a more complete FEA model

that includes a couple-field analysis which investigates the combined effects of the

mechanical electrostatic and capacitive components of the MEMS sensor This allows

for a more accurate prediction of the sensor response as well as the intrinsic sensor

parasitic capacitance

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

64

References

[1 V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[1] V Wowk Machinery Vibration Measurement and Analysis McGraw-Hill Inc 1991

[2] S Nandi H A Toliyat and Xiaodong Li Condition monitoring and fault diagnosis

of electrical motors-a review IEEE Transactions on Energy Conversion vol 20 pp 719-729 2005

[3] S P Beeby G Ensel and M Kraft MEMS Mechanical Sensors Artech House

2004 [4] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998 [5] W J Fleming Overview of automotive sensors IEEE Sensors Journal vol 1 pp

296-308 2001 [6] D S Eddy and D R Sparks Application of MEMS technology in automotive

sensors and actuators Proceedings of the IEEE vol 86 pp 1747-1755 1998 [7] G Brasseur Robust automotive sensors Instrumentation and Measurement

Technology Conference1997 IMTC97 Proceedings of the IEEE Sensing Processing Networking vol 2 pp 1278-1283 vol2 1997

[8] D R Sparjs Packaging of microsystems for harsh environments IEEE

Instrumentation amp Measurement Magazine vol 4 pp 30-33 2001 [9] R G Azevedo Jingchun Zhang D G Jones D R Myers A V Jog B Jamshidi

M B J Wijesundara R Maboudian and A P Pisano Silicon carbide coated MEMS strain sensor for harsh environment applications IEEE 20th International Conference on Micro Electro Mechanical Systems pp 643-646 2007

[10] R R Schoen T G Habetler F Kamran and R G Bartfield Motor bearing

damage detection using stator current monitoring IEEE Transactions on Industry Applications vol 31 pp 1274-1279 1995

[11] R Komanduri and Z B Hou A review of the experimental techniques for the

measurement of heat and temperatures generated in some manufacturing processes and tribology Tribology International vol 34 pp 653-682 10 2001

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

65

[12] H D Bhatt R Vedula S B Desu and G C Fralick Thin film TiCTaC thermocouples Thin Solid Films vol 342 pp 214-220 326 1999

[13] A Datta H Choi and X Li Batch Fabrication and Characterization of Micro-Thin-

Film Thermocouples Embedded in Metal Journal of The Electrochemical Society vol 153 pp H89-H93 2006

[14] B Ando S Baglio N Pitrone N Savalli and C Trigona Bent Beam MEMS

Temperature Sensors for Contactless Measurements in Harsh Environments Instrumentation and Measurement Technology Conference Proceedings IMTC 2008 pp 1930-1934 2008

[15] T J Holroyd Machine condition and dynamic diagnostics via acoustic emission

IEEE Seminar on On-Line Monitoring Techniques for the Off-Shore Industry pp 51-54 1999

[16] S Dolinsek and J Kopac Acoustic emission signals for tool wear identification

Wear vol 225-229 pp 295-303 4 1999 [17] N Tandon and A Choudhury A review of vibration and acoustic measurement

methods for the detection of defects in rolling element bearings Tribology International vol 32 pp 469-480 8 1999

[18] D Ozevin S P Pessiki D W Greve and I J Oppenheim A MEMS transducer for

detection of acoustic emission events IEEE Sensors pp 4 pp 2005 [19] I J Oppenheim D W Greve D Ozevin D R Hay T R Hay S P Pessiki and N

L Tyson Structural tests using a MEMS acoustic emission sensor Smart Structures and Materials 2006 Sensors and Smart Structures Technologies for Civil Mechanical and Aerospace Systems Proceedings of the SPIE vol 6174 pp 26-35 2006

[20] T Toutountzakis C K Tan and D Mba Application of acoustic emission to

seeded gear fault detection NDT amp E International vol 38 pp 27-36 1 2005 [21] V Ferrari A Ghisla D Marioli and A Taroni Silicon resonant accelerometer

with electronic compensation of input-output cross-talk Sensors and Actuators A Physical vol 123-124 pp 258-266 923 2005

[22] S Choa Reliability of MEMS packaging vacuum maintenance and packaging

induced stress Microsystem Technologies vol 11 pp 1187-1196 1030 2005

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

66

[23] S X P Su H S Yang and A M Agogino A resonant accelerometer with two-stage microleverage mechanisms fabricated by SOI-MEMS technology IEEE Sensors Journal vol 5 pp 1214-1223 2005

[24] A A Seshia M Palaniapan T A Roessig R T Howe R W Gooch T R

Schimert and S Montague A vacuum packaged surface micromachined resonant accelerometerJournal of Microelectromechanical Systems vol 11 pp 784-793 2002

[25] S Kon and R Horowitz A High-Resolution MEMS Piezoelectric Strain Sensor for

Structural Vibration Detection IEEE Sensors Journal vol 8 pp 2027-2035 2008 [26] D L DeVoe and A P Pisano Surface micromachined piezoelectric accelerometers

(PiXLs) Journal of Microelectromechanical Systems vol 10 pp 180-186 2001 [27] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Micromachined inertial sensors Proceedings of the IEEE vol 86 pp 1640-1659 1998

[28] C Liu Foundation of MEMS Pearson Prentice Hall 2006 [29] U Krishnamoorthy R H Olsson III G R Bogart M S Baker D W Carr T P

Swiler and P J Clews In-plane MEMS-based nano-g accelerometer with sub-wavelength optical resonant sensor Sensors and Actuators A Physical vol 145-146 pp 283-290 0 2008

[30] B E N Keeler D W Carr J P Sullivan T A Friedmann and J R Wendt

Experimental demonstration of a laterally deformable optical nanoelectromechanicalsystem grating transducer Optics Letters vol 29 pp 1182-1184 0601 2004

[31] CH Liu and T W Kenny A high-precision wide-bandwidth micromachined

tunneling accelerometer Journal of Microelectromechanical Systems vol 10 pp 425-433 2001

[32] T K Bhattacharyya A Ghosh and D Paul Physical modelling of a MEMS based

electron tunneling accelerometer IEEE SAS pp 101-106 2008 [33] W C Tang T H Nguyen and R T Howe Laterally Driven Polysilicon Resonant

Microstructures Sensors and Actuators vol 20 pp 25-32 1115 1989 [34] N Yazdi N Yazdi F Ayazi and K Najafi Micromachined inertial sensors

Proceedings of the IEEE vol 86 pp 1640-1659 1998

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

67

[35] D M Tanner J A Walraven K S Helgesen L W Irwin D L Gregory J R Stake and N F Smith MEMS reliability in a vibration environment 38th Annual International Reliability Physics Symposium Proceedings of IEEE pp 139-145 2000

[36] H Baltes O Brand A Hierlemann D Lange and C Hagleitner CMOS MEMS -

present and future The Fifteenth IEEE International Conference on Micro Electro Mechanical Systems pp 459-466 2002

[37] H Luo G Zhang L R Carley and G K Fedder A post-CMOS micromachined

lateral accelerometer Journal of Microelectromechanical Systems vol 11 pp 188-195 2002

[38] H K Chu J K Mills and W L Cleghorn MEMS-based power disconnect for 42-

V automotive power systems Journal of MicroNanolithography MEMS and MOEMS vol 7 pp 013010 2008

[39] Y Sun S N Fry D P Potasek D J Bell and B J Nelson Characterizing fruit fly

flight behavior using a microforce sensor with a new comb-drive configuration Journal of Microelectromechanical Systems vol 14 pp 4-11 2005

[40] R Legtenberg A W Groeneveld and M Elwenspoek Comb-drive actuators for

large displacements Journal Micromechanics and Microengineering vol 6 pp 320-329 1996

[41] Horsley D A Horsley D A Horowitz R amp Pisano A P (1998)

Microfabricated electrostatic actuators for hard disk drives IEEEASME Transactions on Mechatronics 3(3) 175-183

[42] A P Pisano Resonant-structure micromotors Micro Electro Mechanical Systems

1989 Proceedings an Investigation of Micro Structures Sensors Actuators Machines and Robots IEEE pp 44-48 1989

[43] G K Fedder Simulation of microelectromechanical systems PhD dissertation

University of California Berkeley 1994 [44] J Dong D Mukhopadhyay and P M Ferreira Design fabrication and testing of a

silicon-on-insulator (SOI) MEMS parallel kinematics XY stage Journal of Micromechanics and Microengineering vol 17 pp 1154-1161 2007

[45] S S Rao Mechanical Vibrations 4th ed Prentice Hall 2005 [46] R P van Kampen M J Vellekoop P M Sarro and R F Wolffenbuttel

Application of electrostatic feedback to critical damping of an integrated silicon

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work

68

capacitive accelerometer Sensors and Actuators A Physical vol 43 pp 100-106 5 1994

[47] M Bao H Yang H Yin and S Shen Effects of electrostatic forces generated by

the driving signal on capacitive sensing devices Sensors and Actuators A Physical vol 84 pp 213-219 91 2000

[48] H Baltes O Brand G K Fedder C Hierold J G Korvink O Tabata Enabling

Technologies for MEMS and Nanodevices Wiley-VCH 2004 [49] M Kraft Closed loop digital accelerometer employing oversampling conversion

PhD dissertation Coventry University 1997 [50] M Napoli C Olroyd B Bamieh and K Turner A novel sensing scheme for the

displacement of electrostatically actuated microcantilevers Proceedings of the American Control Conference pp 2475-2480 vol 4 2005

[51] S Franco Design with operational amplifiers and analog integrated circuits 3rd

ed Mcgraw Hill 2002 [52] T B Gabrielson Mechanical-thermal noise in micromachined acoustic and

vibration sensors IEEE Transactions on Electron Devices vol 40 pp 903-909 1993

[53] POLYMUMPS Design Handbook Revision 110 2005 [54] SOIMUMPS Design Handbook Revision 40 2004 [55] M E Motamedi MOEMS SPIE Press 2005 [56] P Hassanpour Design and Analysis of Microelectromechanical Resonant

Structures PhD dissertation University of Toronto 2008 [57] H Luo G K Fedder and L R Carley A 1 mG lateral CMOS-MEMS

accelerometer The Thirteenth Annual International Conference on Micro Electro Mechanical Systems pp 502-507 2000

• Introduction
• • Research Purpose
  • Microelectromechanical Sensors
  • Automotive Sensors
  • Operating Environment
  • Objectives and Contribution
  • Research Outline
  • • Literature Review
    • • Existing Methods for Fault Detection
      • • Motor Current Monitoring
        • Temperature Sensors
        • Acoustic Emission Sensors
        • • MEMS Vibration Sensors
          • • Resonant Sensors
            • Piezoelectric Sensors
            • Displacement Variation Sensors
            • • Sensor Structure
              • • Sensing
                • Mechanical Analysis
                • • Micro-Flexure Selection
                  • Equations of Motion
                  • Quality Factor
                  • • Drive Stability
                    • Electrostatic Actuation
                    • Section Summary
                    • • Capacitance to Voltage Readout Circuit
                      • Noise Analysis
                      • • Mechanical-Thermal Noise
                        • Electrical Noise
                        • • Fabrication Overview
                          • • PolyMUMPS
                            • SoiMUMPS
                            • • Experimental Results
                              • • MEMS sensor ndash T1
                                • Readout Circuitry
                                • MEMS sensorndash L1
                                • Section Summary
                                • • Conclusion and Future Work
                                  • • Conclusion
                                    • Future Work