35 mhz array

224 ieee transactions on ultrasonics, ferroelectrics, and frequency control, vol. 53, no. 1, january 2006

Development of a 35-MHz Piezo-CompositeUltrasound Array for Medical ImagingJonathan M. Cannata, Member, IEEE, Jay A. Williams, Qifa Zhou, Timothy A. Ritter,

and K. Kirk Shung, Fellow, IEEE

AbstractThis paper discusses the development of a 64-element 35-MHz composite ultrasonic array. This array wasdesigned primarily for ocular imaging applications, and fea-tures 2-2 composite elements mechanically diced out of ane-grain high-density Navy Type VI ceramic. Array el-ements were spaced at a 50-micron pitch, interconnectedvia a custom exible circuit and matched to the 50-ohmsystem electronics via a 75-ohm transmission line coaxialcable. Elevation focusing was achieved using a cylindricallyshaped epoxy lens. One functional 64-element array wasfabricated and tested. Bandwidths averaging 55%, 23-dBinsertion loss, and crosstalk less than 24 dB were mea-sured. An image of a tungsten wire target phantom was ac-quired using a synthetic aperture reconstruction algorithm.The results from this imaging test demonstrate resolutionexceeding 50 m axially and 100 m laterally.

I. Introduction

High frequency (> 30 MHz) ultrasound is currentlyused for various imaging applications in ophthalmol-ogy [1][3], dermatology [4], [5], and small animal stud-ies [6], [7]. At present there are several commercial ultra-sound systems available for use in the 25-MHz to 50-MHzfrequency range1. These systems rely upon single-elementtransducers that are mechanically scanned in a line or arcto form an image slice. Array systems, on the other hand,are desired because they use electronic scanning to do so.Arrays also lack movable parts that may be hazardous topatients, can be steered and dynamically focused in theimage plane, and can achieve higher image frame rates.Unfortunately at the present time commercial array sys-tems are not yet available at frequencies above 30 MHz duemainly to limitations in fabrication technology and equip-ment, as well as a lack of quality high-frequency materialsand electronics. Despite these limitations, a few investiga-tors have successfully designed and built high-frequencyultrasonic arrays. The eorts of these researchers are sum-marized next.

Manuscript received April 11, 2005; accepted July 15, 2005. Theauthors would like thank the National Institutes of Health (NIH) forproviding the funding through grant # P41-EB2182.J. M. Cannata, J. A. Williams, Q. Zhou, and K. K. Shung are

with the NIH Resource on Medical Ultrasonic Transducer Technol-ogy, Department of Biomedical Engineering, University of SouthernCalifornia, Los Angeles, CA (e-mail: [email protected]).T. A. Ritter is with the U.S. Air Force, Keesler Air Force Base,

Biloxi, MS.1Systems include those by VisualSonics, Inc., Toronto, Ontario,

Canada, http://www.visualsonics.com; Ultralink LLC., St. Peters-burg, FL, http://www.arcscan.com; and Capistrano Labs, Inc., SanClemente, CA http://www.capolabs.com.

Lower-frequency arrays typically use mechanical dicingto separate array elements, whereby elements are cut froma plate of piezoelectric ceramic or single crystal and back-lled with a polymer ller [8]. This technique generally hasbeen limited to arrays designed to operate in the less-than20-MHz range. However, in recent years, several studieshave proven that it is a viable option for manufacturingarrays with operational center frequencies up to 30 MHz[9][11]. The most sophisticated of the arrays built was the128-element 30-MHz 1-3 piezo-composite array developedby Michau et al. [11]. This array was designed to have anelement-to-element spacing, or pitch, of approximately 2in water or 100 m, with composite posts and array ele-ments both separated using a mechanical dicing saw. Al-ternately, a 48-element 30-MHz array was developed usinga patented process of bonding thin plates of piezoceramicwith a carefully engineered microsphere-loaded polymer tocreate a 2-2 piezo-composite matrix [12], which was thenmechanically diced to form individual array elements [13].This array was also designed for a 100-m pitch. As an al-ternative to mechanical dicing, several investigators havefabricated arrays in the 2030 MHz range using thin sheetsof piezo-polymer materials [14], [15]. Low lateral couplingand low acoustic impedance for polyvinylidene uoride(PVDF) and poly(vinylidene uoride-triuoroethylene)(P(VDF-TrFE)) make these materials good choices forhigh-frequency array design. Unfortunately the low capac-itance of these materials and the high parasitic capaci-tance of typical preamplier inputs preclude the use ofsuitably sized active elements for high-frequency operation[16]. Other keress linear array designs incorporating highdielectric piezoceramics have been evaluated but have yetto be fabricated at high frequencies [17]. For array devel-opment at very high frequencies sol-gel deposition of thinand thick lms of lead-zirconate-titanate (PZT) are viableoptions. Lukacs et al. [18] reported the use of sol-gel depo-sition of PZT lms in the fabrication of 50200MHz single-element transducers and 4060 MHz arrays. Mechanicalseparation of elements was achieved using laser dicing tech-nology. Unfortunately the porosity of the lm producedpoor piezoelectric properties, with a thickness-mode cou-pling coecient (kt) of lower than 25% and a low relativeclamped dielectric permittivity (S33 < 250). Improvementsin this technique are ongoing, with reported piezoelectriccoupling as high as 50% [19]. Sputtering thin PZT lms isanother option for fabrication of very-high-frequency ar-rays, but as with sol-gel deposition this method has beenshown to produce reduced piezoelectric properties when

08853010/$20.00 c 2006 IEEE

cannata et al.: 35-mhz piezo-composite ultrasound array 225

TABLE IInitial Design Goals for the 35-MHz Array.

Center frequency 35 MHzNumber of elements 64Element-to-element spacing (pitch) 50 m (1.2)Elevation aperture 3 mmElevation focus 10 mmBandwidth (6 dB) > 50%20 dB pulse length < 120 nsCrosstalk (element-to-element)


epoxy under vacuum. Another technique that is suitablefor large-scale production is termed the lost mold tech-nique [30]. This technique involves the pressing of a piezo-ceramic paste into a mold, which is usually silicon and canbe designed for various composite geometries. The mold isremoved by chemical etching. After sintering the remain-ing composite posts, the voids left by the mold removal arebacklled with epoxy. It should be noted, however, that atpresent both the tape-cast and lost-mold techniques ex-hibit inferior piezoelectric properties when compared tocomposites made from bulk piezoceramics. Because easeof manufacture and use of bulk piezoceramic plates werecritical design considerations when the 35-MHz array wasdeveloped, the dice and ll method was adopted to man-ufacture a 2-2 composite for this study.

The 35-MHz composite was diced using a 13-mhubbednickel/diamond blade (Asahi Diamond Industrial Co.,Ltd., Tokyo, Japan) and a Tcar 864-1 (Thermocarbon,Inc., Casselberry, FL) dicing saw. The polymer used to llthe kerfs was a mixture of Epo-Tek 301 (Epoxy Technolo-gies, Billerica, MA) and 1-5 m aluminum oxide (Al2O3)particles (Buehler, Ltd., Lake Blu, IL). A plate of NavyType VI ne-grain piezoceramic (TRS600FGHD, TRS Ce-ramics, State College, PA) measuring 10100.5 mm wasselected and waxed to a glass carrier plate using a low-temperature paran wax. The mechanical dicing saw wasrst used to cut kerfs in the ceramic 130 m deep at a 50-m pitch. These rst cuts were subsequently lled with theAl2O3-loaded epoxy (17% Al2O3 by volume) using capil-lary action. A larger volume fraction of Al2O3 was desiredfor this application because it would translate to a highershear wave resonant frequency in the kerf. Unfortunatelya higher percent Al2O3 made the mixture too viscous forthis application. The ller in the rst set of kerfs was leftto cure at room temperature in a dry nitrogen environ-ment for 48 hours. The excess epoxy was then lapped oto expose the diced ceramic. It was important to lap atleast 1020 m into the ceramic to expose the narrowerpart of the kerfs. The dicing saw was then aligned to thecentral portion of the exposed ceramic posts and a secondset of cuts was made to produce the nal 25-m compos-ite pitch. The second set of diced kerfs was subsequentlylled with the same ller as before. After curing, excessepoxy was lapped o the top of the composite. The mostuniform and smallest kerfs were produced at the deepestdicing depths (Fig. 2). Therefore the deepest section of thediced ceramic was chosen to produce the array composite.The composite was lapped to a nal thickness of 50 m af-ter it was placed in the array housing. A scanning electronmicrograph (SEM) image of a composite cross section be-fore placement in the array housing is shown in Fig. 3. Anaverage composite kerf width produced was 14 m with anet piezoceramic volume fraction of 44%.

B. Finite Element Model Optimization of a CompositeArray Element

The nished composite was very fragile in its nal state.It was therefore not tested prior to array fabrication. How-

Fig. 2. An SEM image of the diced ceramic before lapping it to thenal composite thickness. The lighter portions of this image are theceramic posts whereas the darker portions are the kerf ller. Thedashed lines indicate the portion of the diced ceramic used in the35-MHz array.

Fig. 3. An SEM image of a lapped and backed 2-2 composite. Thelighter portions of this image are the ceramic posts whereas thedarker portions are the kerf ller. The irregular textured section un-der the composite is the silver epoxy backing layer.

ever, a nite element model (FEM) was used to predictcomposite and array performance prior to fabrication. Therelevant properties of the active and passive materials usedin the model are listed in Tables II and III, respectively.A 2-D nite element model (PZFLEX, Wiedlinger Asso-ciates, Los Altos, CA) was rst used to generate an elec-trical impedance magnitude and phase plot for a 1.50.3-mm, 50-m-thick virtual piece of fabricated composite res-onating in air (Fig. 4). The series (fS) and parallel (fP )resonance frequencies from this model were used to de-termine the thickness mode coupling coecient kt for thecomposite material based upon the following formula [31]:

kt =

2fSfP

tan[

2

(fP fS

fP

)]. (1)

The average modeled electromechanical coupling coef-cient for the composite was 0.64 and the rst piezoelec-


TABLE IIRelevant Bulk Material Properties for the TRS600FGHD Piezoelectric Ceramic.1

Stiness Dielectric Stress Otherconstants2 constants constants2 properties

cE11 (GPa) 140 S33/

2o 1350 e15 (C/m2) 20.24 Density2 (kg/m3) 7500

cE33 (GPa) 121 S11/

2o 1700 e31 (C/m2) 5.7 V 33 (m/s) 3966

cE12 (GPa) 75 T33/o 3670 e33 (C/m

2) 25.8 k33 0.68cE13 (GPa) 90

T11/o 3830

cE44 (GPa) 22

1Courtesy of TRS Ceramics, State College, PA.2Used in FEM.

TABLE IIIThe Properties of the Passive Materials Used in the Array Design.1

Density Vlong2 Attenuationlong Vshear2 AttenuationshearMaterial (kg/m3) (m/s) (dB/mm) (m/s) (dB/mm)

Epo-Tek 301 1150 2675 13.5 1270 48(Lens)

E-Solder 3022 3200 1850 110 (Backing layer)

Epo-Tek 301 + 17% 1610 2710 15.9 1375 49Al2O3 (Kerf ller)

1All measurements were performed at 30 MHz [32].2Vlong and Vshear are the longitudinal and shear phase velocities, respectively.

trically coupled lateral resonance attributed to the kerfoccurred near 55 MHz. Also, based upon the series reso-nance peak frequency of approximately 41 MHz, the longi-tudinal velocity for the modeled composite geometry was4100 m/s. Therefore, with the ceramic and epoxy densi-ties listed in Tables II and III, respectively, the acousticalimpedance for the modeled composite was calculated tobe approximately 16.7 MRayls.

It has been previously reported that nite element mod-eling can provide an accurate depiction of high-frequencyarray performance [32], as well as reduce the number oftime-consuming prototype fabrication runs. For this studyFEM was used as an aide to determining the optimal ma-terials and geometries used in the array design so that thedesign goals in Table I could be met. For PZFLEX, as wellas any other FEM, the accuracy of the model results arelimited by the accuracy with which the properties of ma-terials used in the design are measured. It would be idealif the complete set of properties were measured over theentire bandwidth of the device. Unfortunately, at high fre-quencies it may be very dicult to characterize some ma-terials due to the increase in attenuation associated withthe elevated frequency. An example of this problem can beseen in Table III. It was not possible to characterize theshear wave velocity and attenuation of the lossy conduc-tive backing material used in the current 35-MHz arraydesign. Therefore, in order to eectively model the arrayusing FEM, a number of assumptions were made to llin these missing material properties. The shear wave at-tenuation was assumed to be four times that of the longi-

tudinal wave attenuation. This approximation was looselybased upon the reported dierence between longitudinaland shear attenuation for other epoxy mixtures measuredat high frequencies [33]. An approximate shear velocitywas calculated based upon the expected Poisson ratio ()and the measured longitudinal velocity using the followingformula [34]:

VshearVlong

=

0.5 1 . (2)

A Poisson ratio of 0.37 was assumed for this material,which is indicative of many rigid epoxies and plastics [34].Therefore the shear velocity was assumed to be 45% thatof the longitudinal velocity reported.

A quarter-wavelength of coaxial cable can provide animproved impedance match between a transducer and theelectronics if it possesses the proper impedance charac-teristics [35]. If properly designed and implemented, thistransmission line coax can serve to increase bandwidthand sensitivity, as well as provide minor adjustments tothe center frequency of the transducer. Cable impedancematching is well suited for use with devices operating athigh frequencies (> 20 MHz), where the length of the coaxcan be approximately three meters or less. For this study, ahigh-impedance (> 50 ohm) micro-coaxial cable was usedto match a high (> 200 at resonance)-impedance ar-ray element to the 50- send/receive electronics. Thistechnique can also be applied to large aperture single-element transducers, which typically display an electrical


Fig. 4. Modeled electrical impedance magnitude (solid line) andphase angle (dashed line) for a 1.5-mm 0.3-mm 0.05-mm pieceof composite resonating in air.

impedance of less than 50 , by using a low impedancecoax [36].

In order to utilize this coax in the nite element modelthe cable was analyzed using the well-known transmis-sion line equations. The rst step was to characterize theimpedance Zo and propagation constant for the coax at35 MHz based upon the following formula for a coax lengthx [37]:

Zx = Zo[Zload + ZoTanh(x)][Zo + ZloadTanh(x)]

, (3)

where Zload represents the electrical impedance of thetransducer and Zx is the transformed coaxial impedancemeasured on the system end of the cable.

The values for Zo and were obtained by measuringthe complex open and short-circuit impedance for a sam-ple cable on an HP 4194 Impedance Analyzer (AgilentTechnologies, Englewood, CO). A given cable length x wasprepared for measurement, and the complex transformedimpedance was recorded for both the short-circuit and theopen-circuit Zload values. The two measured values for Zxand (3) were then used to solve for Zo and . The dis-tributed network representation for the 75- micro-coaxialcable (#171-0574-XX, Precision Interconnect, Portland,OR) used in the array design was also solved [38]. Theresultant per-unit-length values of cable propagation ve-locity (v), series resistance (r), series inductance (l), shuntcapacitance (c), shunt conductance (g), and attenuation() were recorded and listed in Table IV.

Because of the moderately low composite acousticimpedance (16.7 MRayls), only a single matching layerwas used in the array design. Based upon the formula byDeSilets et al. [38] for broadband transducer operation,the ideal matching layer impedance was determined to be3.35MRayls. Therefore the array lens epoxy (Epo-Tek 301,Z = 3.1 MRayls) was chosen and assumed to be a sin-gle /4 matching layer in the 2-D nite element model.From the FEM the ideal matching layer thickness was de-termined to be 19 m and the ideal micro-coaxial cable

TABLE IVMeasured Properties for the 75- Precision Interconnect

Coaxial Cable.1

Property PI #171-0574-XX

Characteristic impedance (Zo) 73.8 1.6i Propagation constant () 0.071 + 0.99i m1Propagation velocity (Vp) 2.2 108 m/sResistance/unit length (r) 6.77 /mCapacitance/unit length (c) 61 pF/mConductance/unit length (g) 674 S/mInductance/unit length (l) 0.33 H/mAttenuation/unit length 0.61 dB/m

1The cable was characterized at 35 MHz. The respective cable pa-rameters are displayed with units of meters (m), ohms (), henrys(H), farads (F), siemens (S), and decibels (dB).

length was 1.2 m. The electrical impedance of a modeledsingle array element before and after adding the coaxialcable is displayed in Fig. 5. The model was nally usedto predict the pulse/echo response from a at plate reec-tor placed at 9.5 mm away from the array face. The echoresponse, shown in Fig. 6, displays a center frequency of34.8 MHz and 6 dB bandwidth of 53%. Once the band-width requirements for the array element in the nite el-ement model were met, the next developmental step wasto fabricate a full 64-element array module as shown pre-viously in Fig. 1.

C. Array Module Fabrication

The conductive backing material, E-solder 3022 (VonRoll Isola, Inc., New Haven, CT), was cast on a lapped andelectroded plate of 2-2 composite material using a 2,000gcentripetal force for ten minutes. This process separatedthe epoxy into loaded and unloaded layers and served tocompact the silver akes up against the composite plate.The backed composite was cured in a dry nitrogen envi-ronment for 24 hours at room temperature and then post-cured at 40C and 50C for 10 hours each. The elevatedtemperature post-cure increased the glass transition tem-perature of the epoxy in order to provide a more rigidsubstrate for the latter array fabrication steps. After thebacking was fully cured, it was lapped down to a thicknessof approximately 3 mm in order to remove the unloadedepoxy layer and to ensure that the silver-loaded portionwas parallel to the composite plate. In this state the back-ing epoxy displayed an electrical resistance of 0.6 /mm.The backed composite was then diced to a width of 3 mmand length of approximately 3.7 mm, to provide at leastve extra elements on either side of the 64-element array.

The backed composite was then cast into a rigid ce-ramic frame (machinable glass-mica ceramic, McMaster-Carr Supply Company, Cleveland, OH) using Epo-Tek 301epoxy. This rigid frame was used to provide support tothe composite, as well as serve as a nonconductive bar-rier surrounding the composite that could be used as asurface for electrical interconnect to individual array ele-ments [13]. The framed composite was allowed to cure at


Fig. 5. Modeled electrical impedance magnitude (solid line) andphase angle (dashed line) for a single composite array element before(top) and after (bottom) coaxial cable impedance matching.

room temperature for 48 hours in a dry nitrogen environ-ment, and then post-cured at 40C and then 50C for 24hours each. The composite and frame were then lappeddown by 10 m in thickness to ensure that the compositeposts were 50-m thick and that the epoxy bond line be-tween composite and frame was at and contiguous. Thetop and side surfaces of the framed composite were thencleaned and sputtered with a total of 4500A Cr/Au. El-ement separation was achieved by mechanically scratch-dicing the top electrode layer over the composite kerfs andceramic frame with the prescribed 50-m spacing using the13-m diamond/nickel hubbed blade previously described.The opposing two electroplated sides of the ceramic framewere also scratched-diced to separate individual elementelectrodes. For this a 50-m blade was used to completelyremove a 50-m width of electroplating, leaving connec-tions for either the odd or even element electrodes on oneside of the ceramic frame spaced at a 100-m pitch (Fig. 1).The electroplated array was then prepared for casting ofthe epoxy lens.

The epoxy lens was cast onto the array with a polishedquartz cylindrical rod (ISP Optics, Irvington, NY) as amold. The desired focal length f was achieved by specify-

Fig. 6. The FEM pulse-echo response for a single 35-MHz arrayelement. The axes on the top and right of the gure refer to thefrequency spectrum (dashed line). Also displayed are the calculatedcenter frequency (CF), the 6 dB bandwidth (BW), and the 20 dBpulse length (PL) for the response.

ing the radius of curvature of the lens using the followingformula [39]:

f 1 c2

c1

, (4)

where is the radius of curvature of the lens, c1 is thesound velocity in the lens material, and c2 is the soundvelocity in the medium. A quartz rod with a radius ap-proximately equal to 4.34 mm was chosen to focus thearray at a range of 10 mm. The quartz rod was sprayedwith a mold release agent (Ease Release 200, Mann For-mulated Prods., Gillete, NJ) to ease in the epoxy post-cureremoval process. As noted previously, the apex of the lenswas designed to be approximately /4 thick at the centerof the elevation aperture to aid in acoustically matchingthe transducer to the load medium. To ensure this, thecylindrical rod was centered over the elevation apertureand oset by 19 m using shims. The lens epoxy, Epo-Tek 301, was degassed and ltered to remove bubbles andimpurities prior to application. A bolus of the epoxy wasallowed to wick between the cylindrical rod and array as-sembly. Surface tension held the liquid epoxy in place dur-ing curing. The lens was allowed to cure for two days in adry nitrogen environment before post-curing at 40C for 4hours. After curing, the edges of the lens above the ceramicframe were lapped parallel to the array elements.

D. Array Interconnect, Housing, and Termination

The connection of individual coaxial cables to arrayelements was accomplished through the use of two in-


Fig. 7. A picture of the nished 35-MHz array. The bottom right sideof this image shows an enlarged view of the front of the array. Thelong axis of the hole cut in the stainless steel nosepiece is along theazimuth direction.

termediate exible circuits incorporating 5-m-thick Autraces on a 50-m-thick polyimide surface (Dynamic Re-search Corp., Wilmington, MA). The coaxial cable as-sembly (#171-0574-XX, Precision Interconnect, Portland,OR) was attached to the exible circuit by the cable man-ufacturer using a proprietary low-temperature solder pro-cess. The exible circuit was connected to the framed arrayby carefully aligning and bonding the ex circuit tracesto the electrodes on the two sides of the framed array.The epoxy used for this process was Epo-Tek 301. Thenewly interconnected array assembly was allowed to curefor 48 hours at room temperature in a dry nitrogen envi-ronment before further processing. A grounding wire wasconnected from the conductive epoxy array backing to thecable assembly shield. The array assembly was housed in ave-part cylindrical metal and plastic housing. Individualcoaxial cables were terminated into an aluminum enclo-sure housing 64 individual SMB connectors. This ensuredsatisfactory electrical isolation between each coaxial cable.A photograph of the completed array is shown in Fig. 7.

E. Array Characterization

The ultimate indication of ultrasonic array performanceis its ability to form an image. However, image quality de-pends not only on the array but also on the system elec-tronics and signal processing algorithms. Thus it is dicultto compare dierent arrays based solely on image anal-ysis. Therefore, several standard non-imaging tests wererst performed on the 35-MHz array so that it could becompared to arrays built prior to this study.

The electrical impedance magnitude and phase angle foreach element were measured using a HP4194 impedanceanalyzer with the z-probe attachment. For this test thearray face was placed in a water bath, and the electricalimpedance magnitude and phase angle were measured overthe array pass-band.

The pulse-echo response test is the most common testperformed on ultrasonic single-element and array trans-ducers. This test was used to measure and provide a rela-tive comparison of array element center frequency, band-width, focus, pulse length, and sensitivity. A Panamet-rics 5900PR pulser/receiver (Panametrics, Inc., Waltham,MA) was used to excite each element using the 1-J, 50-settings with 20 dB total gain on receive. The array waspositioned in a degassed/deionized water bath opposite apolished stainless steel plate reector. The time delay ob-served on the oscilloscope (Lecroy LC534, Chestnut Ridge,NY) between excitation and the largest rst echo responsewas recorded to determine elevation focal length. The fastFourier transform (FFT) function on the oscilloscope wasused to determine the frequency response of this RF echo.The two 6 dB points of this power spectrum dened theupper and lower band edges of the signal, and the meanof these frequencies was recorded as the center frequency.The percent bandwidth was calculated by dividing the fre-quency dierence between the two 6 dB points by thecenter frequency. The amplitude of the echo signal wasrecorded for relative element sensitivity comparisons. Thepulse length of the echo waveform was recorded as thelength of time between the rst and last points where thesignal was 20 dB relative to the peak.

The level of electrical and acoustical separation betweenelements was determined by measuring crosstalk. For thistest the array was positioned in a degassed/deionizedwater bath opposite an absorptive piece of rubber. ASony/Tektronix AFG2020 function generator (Tektronix,Inc., Beaverton, OR), set in burst mode, was used to ex-cite a representative element with the applied voltage mea-sured as a reference. Voltages across the three nearest ele-ments were measured and compared to this reference volt-age. This process was repeated at discrete frequencies overthe array pass-band.

The azimuthal one-way directivity was measured by ro-tating a representative element around an axis along itslength and center. A needle hydrophone (Precision Acous-tics, Dorchester, UK), placed at the elevation focus, wasused to acquire the amplitude of the time-domain responseat discrete angular positions. The Field II program wasused to simulate the directivity of a single array element[40], and to estimate the eective element width by match-ing a theoretical directivity curve to the measured values.

Insertion loss was measured by exciting a representa-tive array element with a 35-MHz voltage burst, and re-ceiving the reected echo from a polished steel reectorplaced at the elevation focus. The receive power across a50- load was referenced to the source power delivered toa 50- reference load and expressed in decibels. The mea-sured value was then corrected for loss due to diraction inthe azimuth direction [41], attenuation in the water bath(2.2 104 dB/mm-MHz2) [42], and reection from thesteel target (0.6 dB).

For the nal performance test, the array was used toimage a custom-made wire target phantom composed ofve evenly spaced 20-m-diameter tungsten wires (Cali-


fornia Fine Wire Company, Grover Beach, CA). There arecurrently two 16-channel analog [43] and digital [44] high-frequency beamformers available to produce real-time im-ages at center frequencies up to 40 MHz. Unfortunately,given the current geometry of the 35-MHz array, it was de-termined that 16 transmit and receive channels would notprovide adequate lateral resolution for this study. There-fore a synthetic aperture reconstruction algorithm wasused to form the image of the wire target phantom.

The 1.2- pitch is problematic for synthetic apertureimaging because the rst-order grating lobe appears atone-half the angular location observed in a conventionalbeamforming system [45]. For the 35-MHz array withoutbeam steering, grating lobes are expected to occur at 57

and 28.5 for conventional and synthetic systems, respec-tively. To limit the eect of these secondary lobes, ar-ray elements were only used for reconstruction if they fellwithin a 9 acceptance angle for a point in the formedimage. Therefore, the aperture size used for reconstruc-tion was varied dynamically throughout the image to pro-duce a consistent lateral resolution. Backprojection usinga monostatic approach was accomplished by delaying andsumming the time-domain contributions to each pixel ac-cording to the following formula [13]:

P (xi, zi) =N

e=1

weRe

[t 2

c

(xe xo)2 + z2o],(5)

where xi, zi is the location of the pixel in the image plane,e is the index of the element over the range of 1 to N(number of elements), we is the apodization function, Reis the time-domain response, t is the time, c is the propa-gation velocity, xo, zo is the location of the point in objectspace, and xe is the position of the array element. The indi-vidual time-domain responses were acquired by manuallyconnecting each of the 64 array elements to a Panametrics5900 pulser-receiver using the same settings as describedfor the pulse-echo test setup. Image reconstruction was ac-complished o-line using programs written in Matlab [25].

III. Results and Discussion

All of the individual element test results for the com-pleted array are summarized in Table V. The array dis-played only one open element and no shorted elements.Fig. 8 shows the electrical impedance magnitude andphase angle for a typical array element. The general trendof this curve matches that of the FEM impedance plotshown in Fig. 5. However, the resonance peak seen in theFEM plot near 40 MHz, which corresponds to the piezo-composite parallel resonance, is not visible in the measuredimpedance plot. The likely cause of this discrepancy is ahigher-than-desired electrical trace resistance between anarray element and a coaxial cable. Exploring the FEM fur-ther veried that increasing element trace resistance wouldproduce the measured result. According to the model, the

TABLE VMeasured Properties for the 64-Element Composite Array.

Property Value

Number of elements 64Number of open elements 1Number of shorted elements 0Average center frequency 35.3 MHzHighest/lowest center frequency 36.5 MHz/34.2 MHzAverage bandwidth (6 dB) 55%Highest/lowest bandwidth 62%/49%Average sensitivity 403 mVHighest/lowest sensitivity 444 mV/360 mV20 dB pulse length 94 nsElectrical impedance magnitude

(at 35 MHz) 31 ohmsElectrical impedance phase angle

(at 35 MHz) 48 degreesFocal point 9.5 mmInsertion loss 22.8 dB

Measured on a representative array element.

Fig. 8. Measured electrical impedance magnitude (solid line) andphase angle (dashed line) for a representative composite array ele-ment.

consequence of this is a slight reduction of pulse sensitivityand bandwidth.

A comparison of the pulse-echo responses for a typicaland the worst (subjectively judged by the authors) con-nected array element is shown in Fig. 9. The array dis-played an average center frequency of 35.3 MHz with anaverage 6 dB bandwidth of 55%. Both of these aver-age measurements were consistent with the initial require-ments for the array design and the nite element model.The measured 20 dB pulse length for a typical elementof 94 ns also met the initial specications for the array.However, this recorded pulse length is slightly longer thanexpected from the FEM. This error between model andexperiment can be attributed to the observed additionalresistance in the element traces, as well as the expectederrors due to the aforementioned assumptions used in themodel. A larger increase in positive trace resistance is alsothe most likely cause of the additional reduction of sensi-


Fig. 9. Measured pulse-echo response for two 35-MHz array elements.The waveform on the top (element #64) was deemed typical whereasthe one on the bottom (element #24) was judged to be the worst.The axes on the top and right of the gure refer to the frequencyspectrum (dashed line). Also displayed are the calculated center fre-quency (CF), the 6 dB bandwidth (BW), the 20 dB pulse length(PL), and peak-to-peak voltage (Vpp).

tivity and pulse shape recorded for the worst active ele-ment. Overall the performance of the array was very con-sistent across the aperture, as shown in Figs. 10 and 11.

Crosstalk measurements indicated satisfactory but notideal element-to-element isolation. The measured crosstalkfor the array is shown in Fig. 12. Near the center frequencyof the array the measured crosstalk was


Fig. 13. Measured and modeled one-way directivity for a single-arrayelement. This gure shows three sets of hydrophone measurementstaken at a range of 10 mm, and the modeled result also taken at arange of 10 mm using a 70-m element width.

element impedance and the impedance between successiveelectrodes. Although a combination of the two sources islikely, coupling was observed to be instantaneous, indicat-ing that electrical crosstalk was dominant. As a compari-son, Guess et al. [46] observed combined crosstalk levelsexceeding 24 dB for adjacent elements on a similarlyconstructed 3.0-MHz array. However, that study showedthat, while high crosstalk levels were observed for near-est neighbors, crosstalk between next nearest neighbor el-ements was much lower at less than 45 dB over the pass-band. This was not the case for the 35-MHz array design.In fact, the crosstalk magnitude for the next nearest neigh-bor element, which had a ex circuit connection next tothe source element, was comparable to the magnitude ob-served for the nearest neighbor. This observation furtherreinforces the notion that the major source of crosstalk inthe fabricated array was electrical in nature. In the futurethis issue could be overcome by improving the electricalimpedance characteristics of an array element via reducedelectrode trace resistance, and/or by increasing elementarea or dielectric permittivity of the piezo-ceramic used inthe design.

The measured one-way directivity pattern for a rep-resentative array element is shown in Fig. 13. Threedata sets were taken for comparison. The higher-than-desired crosstalk observed explains why the eective ele-ment width (70 m) was 40% larger than the array pitch.A similar result was observed by Felix et al. for a 3.5-MHzarray [47]. The null in the peak of these patterns can alsobe attributed to crosstalk eects [41].

The measured insertion loss for a representative arrayelement was 50.3 dB. Correction for diraction in the az-imuth direction was achieved using the directivity patternfor a 70-m element. After compensation for attenuationcaused by the water and reection from the target, anestimated insertion loss of 22.8 dB was obtained. This

Fig. 14. Synthetic aperture image of a wire phantom reconstructedusing a half-angle of 9 and no apodization or thresholding. Thedynamic range of the image is 40 dB and the display uses a lineargray scale for mapping.

value was comparable to a similarly built, high-frequencylens-focused transducer [36], but was 7.8 dB higher thanreported by Ritter et al. [13] for an optimized 30-MHz 2-2 composite array with two acoustic matching layers andlens. However, given the simplicity of the 35-MHz arraydesign, lower element sensitivity is a compromise in thismore mass-producible design.

An image of the ve-wire test phantom is shown inFig. 14 using a linear gray scale and 40-dB dynamic range.No apodization or thresholding was implemented duringreconstruction. Plots of the axial and lateral line spreadfunctions for the center wire are shown in Fig. 15. Themeasured FWHM (full-width half-maximum) resolutionswere 42 m and 95 m for the axial and lateral directions,respectively. These measurements correlate well with theresolutions of 40 m axially and 95 m laterally predictedfrom a Field II simulation and oer a slight improvement


Fig. 15. Axial (top) and lateral (bottom) line spread functions forthe center wire of the phantom.

laterally and signicant improvement axially over the res-olutions reported by Ritter et al. [13] for the 30-MHz com-posite array. The phantom image displayed low-level side-lobes resulting from the reconstruction algorithm and thelimited number of elements used for reconstruction. Thesesidelobes could be reduced in amplitude by using apodiza-tion; however, this comes at a cost of increased main lobewidth [48].

IV. Conclusions

This paper describes the development and fabricationof a 35-MHz linear array made using mechanically diced2-2 composite elements. With the exception of the desired

50% bandwidth,and a pulse length of less than 120 ns. A wire phantom im-age, reconstructed using a synthetic aperture technique,demonstrated resolutions of 95 m or less with minimalartifacts. The results indicate that this array design, whencoupled with an adequate high-frequency beamformer, willbe suitable for clinical applications. Assessment of the ar-ray on animals and human patients is anticipated in thenear future.

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Jonathan M. Cannata (S01M04) wasborn in Carson, CA on August 4, 1975. Hereceived his B.S. degree in Bioengineeringfrom the University of California at San Diegoin 1998, and his M.S. and Ph.D. degrees inBioengineering from the Pennsylvania StateUniversity, University Park, PA, in 2000 and2004, respectively.

Since 2001 Dr. Cannata has served as theresource manager for the NIH (National In-stitutes of Health) Resource Center for Medi-cal Ultrasonic Transducer Technology at Penn

State University (2001 to 2002) and currently at USC (University ofSouthern California). In 2005 he was appointed the title of ResearchAssistant Professor of Biomedical Engineering at USC. His currentresearch interests include the design, modeling, and fabrication ofhigh frequency single element ultrasonic transducers and transducerarrays for medical imaging applications. Dr. Cannata is a member ofthe Institute of Electrical and Electronics Engineers (IEEE).

Jay A. Williams was born on June 14,1955, in Bellefonte, PA. He graduated fromState College Area School District of PA in1973 and audited courses in Physics, Com-puter Science, Mechanical Engineering, Ar-chitecture, and Business Management at thePennsylvania State University, 19751977. Hehad been in industry for over twenty-ve yearsin technical elds such as Mass Spectrometry,Microwave Telecommunication, Liquid Chro-matography, Digital Electronics, Ultrasound,and Information Technology. From 1990 until

2001, he worked in ultrasonics at Blatek, Inc., eight-and-a-half yearsin engineering and two-and-a-half years in management, includingve years as Network Administrator and two years as Quality SystemManager in charge of establishing their ISO9001/CGMP compliantquality system.


Mr. Williams joined the NIH Resource Center for Medical Ultra-sonic Transducer Technology in March, 2002 as Research Technolo-gist, just prior to their move to the University of Southern Californiain August of that year. He has also served as the webmaster for theResource Center website http://bme.usc.edu/UTRC since the moveto Los Angeles, CA. He has now worked in the ultrasound eld forover fourteen years.

His research interests are in novel techniques for high-frequencytransducer/array design and fabrication, and the role informationtechnology can play in improving the capabilities and accessibility oftechnology.

Qifa Zhou received his Ph.D. degree from theDepartment of Electronic Materials and En-gineering at Xian Jiaotong University, Xian,China in 1993.

He is currently a Research Assistant Pro-fessor at the NIH Resource on Medical Ul-trasonic Transducer Technology (Los Angeles,CA) and the Department of Biomedical En-gineering at University of Southern Califor-nia, Los Angeles, CA. Before joining USC in2002, he worked in the Department of Physicsat Zhongshan University of China at Guang

Zhou, the Department of Applied Physics at Hong Kong Polytech-nic University at Kowloon, and the Materials Research Laboratoryat Pennsylvania State University, University Park, PA.

His current research interests include the development of ferro-electric thin lms, MEMS technology, modeling and fabrication ofhigh frequency ultrasound transducers and arrays for medical imag-ing applications. He has published more than 70 papers in this area.He is a member of the Materials Research Society.

Timothy Ritter was born in Harrisburg, PAon February 19, 1965. He earned his B.S. de-gree in Mechanical Engineering from PennState University in 1987, his M.S. degree inPhysics from the University of Connecticutin 1991, and his Ph.D. degree in Bioengineer-ing from Penn State University in 2000. From1998 until 2001 he served as manager of theNIH Resource Center for Medical UltrasonicTransducer Technology. In 2000 he was alsoappointed an Assistant Professor of Bioengi-neering at Penn State. He accepted a commis-

sion in the U.S. Air Force in 2001 and was assigned to Keesler AirForce Base following completion of Ocer Training School. His cur-rent assignment is in therapeutic and diagnostic medical physics. Heis a member of the American Association of Physicists in Medicineand the Health Physics Society.

K. Kirk Shung (S73M75SM89F93)obtained a B.S. degree in electrical engineer-ing from Cheng-Kung University in Taiwanat Tainan in 1968, a M.S. degree in electri-cal engineering from University of Missouri,Columbia, MO in 1970 and a Ph.D. degreein electrical engineering from University ofWashington, Seattle, WA, in 1975. He didpostdoctoral research at Providence MedicalCenter in Seattle, WA, for one year before be-ing appointed a research bioengineer holdinga joint appointment at the Institute of Ap-

plied Physiology and medicine, Seattle, WA. He became an assistantprofessor at the Bioengineering Program, Pennsylvania State Univer-sity, University Park, PA in 1979 was promoted to professor in 1989.He was a Distinguished Professor of Bioengineering at Penn Stateuntil September 1, 2002 when he joined the Department of Biomedi-cal Engineering, University of Southern California, Los Angeles, CA,as a professor. He has been the director of NIH Resource on MedicalUltrasonic Transducer Technology since 1997.

Dr. Shung is a fellow of the IEEE, the Acoustical Society of Amer-ica and the American Institute of Ultrasound in Medicine. He is afounding fellow of the American Institute of Medical and BiologicalEngineering. He has served for two terms as a member of the NIHDiagnostic Radiology Study Section. He received the IEEE Engineer-ing in Medicine and Biology Society early career award in 1985 andcoauthored a best paper published in IEEE Transactions on UFFC in2000. He is the distinguished lecturer for the IEEE UFFC society for20022003. He was elected an outstanding alumnus of Cheng-KungUniversity in 2001.

Dr. Shung has published more than 160 papers and book chap-ters. He is the author of a textbook Principles of Medical Imagingpublished by Academic Press in 1992. He co-edited a book Ultra-sonic Scattering by Biological Tissues published by CRC Press in1993. Dr. Shungs research interest is in ultrasonic transducers, highfrequency ultrasonic imaging, and ultrasonic scattering in tissues.

35 mhz array

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