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NORTHEASTERN UNIVERSITY

Hybrid Bimetallic-Thermite Reactive Composites:

Ultrasonic Powder Consolidation, Ignition

Characterization and Application to Soldering

A Dissertation Presented

By

Somayeh Gheybi Hashemabad

to

The Department of Mechanical and Industrial Engineering

in partial fulfillment of the requirements

for the degree of

Doctor of Philosophy

in the field of

Mechanical Engineering

Northeastern University

Boston, Massachusetts

December, 2015

i

Acknowledgments

I would like to thank my adviser Professor Teiichi Ando for his guidance and constant

support during my doctoral studies. I consider myself very fortunate to have the opportunity to

work with him. I would also like to thank my dissertation committee members, Professors Yung

Joon Jung and Zhiyong Gu for their invaluable advice and suggestions. Many thanks to

Professors Peter Y Wang, Julie Chen, Charalabos Doumanidis, Zhiyong Gu, Claus Rebholz and

their research teams for our strong collaboration and useful discussions.

I would like to express my sincere appreciation to all my colleagues, Dr. Dinc Erdeniz,

Nazanin Mokaram, Azin Houshmand, Mina Yaghmazadeh, Ming L Wood, Gianmarco Vella,

Tianyu Hu, Zheng Liu, Yangfan Li and Mingze Chen at the Advanced Materials Processing

Laboratory for their help and friendship. Many thanks to Hamid Ebrahimi for his friendship and

help with the tensile tests. I am also thankful to the faculty and staff of the Department of

Mechanical and Industrial Engineering at Northeastern University.

This work made use of CMSE Shared Experimental Facilities at MIT, supported by the

National Science Foundation under award number CMMI-1029758. I would like to acknowledge

Dr. Charlie Settens, Dr. Scott Speakman and William Fowle for their help with material

characterization. Special thanks to Fukuda Metal Foil & powder Co., Ltd. for the supply of Al

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and Ni flakes and technical staff of STAPLA Ultrasonics Corporation, Wilmington, MA for

calibrating the ultrasonic welding unit.

This research was part of a joint project at Northeastern University, University of

Massachusetts-Lowell and Tufts University funded by the National Science Foundation under

award number CMMI-1029758.

Finally, I would like to thank my parents, Narges Ghofrani and Ali Gheybi, my family

and friends for their constant encouragement and support during my education.

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Abstract

Nanoheaters reactive composites have constituents in nano-scale in at least one

dimension and when ignited can generate heat in calculated amounts through self-propagating

exothermic reactions among the constituents. Due to their metastable nature, fabrication of

nanoheaters requires a method that does not involve exposure to high temperature. Ultrasonic

powder consolidation (UPC) is one such process in which powder particles are consolidated

under the action of ultrasonic vibration at low to moderate temperature for a short duration of

time, usually less than a few seconds. In this study, nano-thick flakes of aluminum and nickel

prepared by a proprietary ball milling process and thermal plasma-synthesized nanoparticles of

Fe2O3 and CuO were successfully consolidated by UPC into bimetallic Al-Ni nanoheaters,

thermite Al- Fe2O3 and Al-CuO neanoheaters and new Al-Fe2O3-x(Al-Ni) and Al-CuO-x(Al-Ni)

hybrid bimetallic-thermite nanoheaters. Full-density consolidation with metallurgical bonding

through the Al nanoflakes was achieved for all nanoheaters at consolidation temperatures as low

as 573 K in just 1 s.

Consolidated Al-Fe2O3-x(Al-Ni) and Al-CuO-x(Al-Ni) hybrid bimetallic-thermite

nanoheaters ignited upon continuous heating at or below the melting point of Al. Thermal

analysis showed that the ignition was preceded by an acceleration period in which solid-state

reactions between Al and Ni flakes caused the nanoheater to self-heat. During the acceleration

iv

period, hot spots were created in the nanoheater where the Al-Al3Ni eutectic liquid eventually

formed at 913 K and triggered the thermite reaction which propagated rapidly throughout the

nanoheater. Completion of the exothermic reactions in the nanoheaters was confirmed by X-ray

diffraction which identified the expected final product phases and nothing else.

Combining the excellent ignitability of bimetallic Al-Ni nanoheaters and the high heat

output of thermites, Al-Fe2O3-x(Al-Ni) and Al-CuO-x(Al-Ni) hybrid bimetallic-thermite

nanoheaters are well-suited for use as a heat source in micro-joining by various methods such as

fusion welding, brazing and soldering. In this study, they were tested as heat sources for the

soldering of aluminum and copper. To this end, effective flux-less soldering of aluminum was

also realized in a novel process that cleverly combines ultrasonic abrasive activation of

aluminum surface and subsequent reflow for which a nanoheater can supply heat in a controlled

amount. The flux-less soldering process has been proven to produce metallurgically sound solder

joints. Aluminum 1100 plates joined with SN100C®

solder showed joint strengths of 45 - 48

MPa, values expected for the Pb-free solder.

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Table of content

ACKNOWLEDGMENTS ............................................................................................................. I

ABSTRACT ................................................................................................................................. III

TABLE OF CONTENT ................................................................................................................ V

LIST OF FIGURES ................................................................................................................. VIII

LIST OF TABLES ................................................................................................................. XVIII

1. INTRODUCTION ................................................................................................................. 1

2. BACKGROUND ................................................................................................................... 4

2.1 REACTIVE COMPOSITES AND THEIR APPLICATION IN JOINING .......................................... 4

2.2 IGNITION METHOD .............................................................................................................. 6

2.3 ALUMINUM SOLDERING ...................................................................................................... 8

3. ULTRASONIC CONSOLIDATION OF REACTIVE COMPOSITES ......................... 11

3.1 EXPERIMENTAL PROCEDURE ............................................................................................ 12

3.1.1 Materials ................................................................................................................... 12

3.1.2 Experimental Setup .................................................................................................. 16

3.1.3 Continuous ignition test ........................................................................................... 21

3.2 RESULTS AND DISCUSSION ............................................................................................... 22

3.2.1 Mixing Process ......................................................................................................... 22

3.2.2 Microstructure .......................................................................................................... 25

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3.2.3 Ignition Test ............................................................................................................. 29

4. EARLY-STAGE REACTION KINETICS MODELING ............................................... 47

4.1 THEORY ............................................................................................................................. 48

4.1.1 Total Volume of Product Phase ............................................................................... 48

4.1.2 Heat Generation ....................................................................................................... 51

4.1.3 Thermal Balance ...................................................................................................... 52

4.1.4 Numerical Computation ........................................................................................... 54

4.2 RESULTS AND DISCUSSION ............................................................................................... 60

5. IGNITION METHODS FOR POSSIBLE INDUSTRIAL APPLICATIONS ............... 65

5.1 SPARK IGNITION ................................................................................................................ 66

5.1.1 Experimental Setup .................................................................................................. 66

5.1.2 Results and Discussions ........................................................................................... 68

5.2 JOULE HEATING IGNITION ................................................................................................. 70

5.2.1 Experimental Procedure ........................................................................................... 70

5.2.2 Results and Discussion ............................................................................................. 73

5.3 MICROWAVE HEATING IGNITION ...................................................................................... 78



6. MICROSCALE JOINING ................................................................................................. 87

6.1 COPPER SOLDERING USING REACTIVE COMPOSITES AS HEAT SOURCES ......................... 87

6.1.1 Preparation of Solder Sheet ...................................................................................... 88


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6.2 FLUX-LESS SOLDERING OF ALUMINUM ............................................................................ 93


6.2.2 Results and Discussion ........................................................................................... 100

6.2.3 Aluminum Flux-Less Soldering Using Reactive Composites ................................ 111

7. CONCLUSIONS ............................................................................................................... 114

8. RECOMMENDED FUTURE WORK ............................................................................ 117

REFERENCES ........................................................................................................................... 120

APPENDIX: MATLAB CODES .............................................................................................. 133

viii

List of Figures

Figure 3.1: SEM pictures of as-received flakes (a) Ni and (b) Al. ........................................... 13

Figure 3.2: SEM pictures of as-received oxide porticles: (a) Fe2O3 nanoparticles, (b) CuO

nanoparticles and (c) Al2O3 particles. .................................................................... 14

Figure 3.3: Schematic view of an ultrasonic welder modified for using consolidates powder by

die and punch. ......................................................................................................... 16

Figure 3.4: Schematic diagrams showing the sonotrode tip. .................................................... 17

Figure 3.5: A schematic view of the heater plate...................................................................... 18

Figure 3.6: The control box for the heater plate. ...................................................................... 19

Figure 3.7: Wiring diagram for the heater control box. ............................................................ 19

ix

Figure 3.8: Schematic representation of the continuous heating ignition test setup. ................ 22

Figure 3.9: Effect of mixing method on distribution of particles, particles mixed with (a)

method I, (b) method II, (c) method III and (d) method IV. .................................. 24

Figure 3.10: SEM image of an (a) Al-Ni (b) Al-CuO reactive composite consolidated for 1 s at

523 K under 100 MPa uniaxial pressure. ............................................................... 26

Figure 3.11: SEM images of (a) 2Al-Fe2O3-3(Al-Ni), (b) 2Al-3CuO-1(Al-Ni) and (c) 2Al-

Al2O3-3(Al-Ni) reactive composites, consolidated by subjecting premixed

compacts to ultrasonic vibration for 1 s at 573 K under 100 MPa uniaxial pressure.

................................................................................................................................ 28

Figure 3.12: XRD patterns of ignited hybrid composites. (a) 2Al-Fe2O3-3(Al-Ni) and (b) 2Al-

3CuO-3(Al-Ni) ....................................................................................................... 30

Figure 3.13: The change in specimen temperature with time of a 2Al-3CuO-3(Al-Ni) composite

consolidated at 573 K for 1 s under 100 MPa (a) acceleration stage above 855 K

followed by melting of aluminum and ignition at 895 K, (b) Tsp vs. t near/at

ignition. ................................................................................................................... 31

x

Figure 3.14: SEM image of a 2Al-Al2O3-3(Al-Ni) composite continuously heated and held at

773 K for 5 minutes. Al3Ni formed between Ni flakes and Al matrix. ................. 32

Figure 3.15: Schematic illustration of liquid formation (a) at the reaction front and (b) its effect

on the onset of thermite reaction. ........................................................................... 34

Figure 3.16: The change in specimen temperature with time of a 2Al-Fe2O3-1(Al-Ni) composite


followed by melting of aluminum and ignition at 933 K, (b) Tsp vs. t near/at

ignition. ................................................................................................................... 36

Figure 3.17: The change in specimen temperature with time for 2Al-Al2O3-3(Al-Ni) composite


followed by melting of aluminum and ignition at 829 K, (b) shallow endothermic

dip just before ignition. .......................................................................................... 38

Figure 3.18: Dependence of (a) the ignition and (b) acceleration temperatures of 2Al-Fe2O3-

x(Al-Ni), 2Al-3CuO-x(Al-Ni) and 2Al-Al2O3-x(Al-Ni) composites on x

(bimetallic Al-Ni addition) ..................................................................................... 41

xi

Figure 3.19: The height of the spike before ignition depended on x, the bimetallic Al-Ni

addition: (a) no spike appeared at x = 1, (b), (c) spike appeared at x = 2 and 3 with

its height increasing with increasing x. .................................................................. 42

Figure 3.20: Specimen temperature vs. time of 2Al-Al2O3-x(Al-Ni) composites. No spike nor

sharp endothermic shift preceded ignition, reflecting the lower rates of Al-Ni

reactions. Modified time (t’) is used to place the three thermographs in one

figure. ..................................................................................................................... 43

Figure 3.21: High effectiveness of Fe2O3 and CuO nanoparticle clusters in retarding the

acceleration stage over that of Al2O3 particles. ..................................................... 44

Figure 3.22: Ignition temperature and energy output of (a) 2Al-Fe2O3-x(Al-Ni) and (b) 2Al-

CuO-x(Al-Ni) against x. The energy output is calculated for a unit volume of

composites assuming completion of the reactions in Table 3.1. Literature value

(1233 K [87]) is used for x = 0 since 2Al-Fe2O3 specimens consolidated by UPC

did not ignite at the maximum temperature reachable with the experimental setup.

................................................................................................................................ 46

Figure 4.1: A schematic showing compound (α) particles forming at the A-B interface. ........ 49

xii

Figure 4.2: Compound disks impinging on each other only in lateral growth. ........................ 51

Figure 4.3: Cross section of a 2Al-Fe2O3-3(Al‒Ni) reactive composite ultrasonically

consolidated at 573 K under 100 MPa for 1 s. ....................................................... 55

Figure 4.4: (a) Ni particles in specimen (b) Ni boundaries in the 2Al-Fe2O3-3(Al-Ni)

specimen. ................................................................................................................ 56

Figure 4.5: Al particles in specimen (b) Al boundaries in the 2Al-Fe2O3-3(Al-Ni) specimen. 56

Figure 4.6: The Al-Ni interfacial length. .................................................................................. 57

Figure 4.7: Considered layers in a specimen. ........................................................................... 58

Figure 4.8: Temperature change in a 2Al-Fe2O3-3(Al-Ni) specimen while being held at 500 K,

for nonadiabatic condition. ..................................................................................... 61


500 K and 550 K, for adiabatic condition. ............................................................. 62

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500 K and 550 K, for non-adiabatic condition. ...................................................... 63

Figure 5.1: A fine-wire welder used for spark ignition tests. ................................................... 67

Figure 5.2: Different stages of spark ignition of a 2Al-3CuO-3(Al-Ni) specimen consolidated

by subjecting premixed compacts to ultrasonic vibration for 1 s at 573 K under 100

MPa uniaxial pressure, (a) before applying spark, (b) application of spark and (c)

ignition products. .................................................................................................... 68

Figure 5.3: XRD pattern of 2Al-3CuO-1(Al-Ni) ignited by spark ignition method. ................ 69

Figure 5.4: (a) Schematic of consolidated specimen with two buried copper stripes. (b) A

possible path of electrons that pass through two copper strips. ............................. 70

Figure 5.5: Schematic of resistance measurement .................................................................... 72

Figure 5.6: Schematic of resistance measurement method. ...................................................... 73

xiv

Figure 5.7: Reduction of resistance by improving the mixing.................................................. 75

Figure 5.8: Example image of the gap measurement method using the optical microscope

image. ..................................................................................................................... 76

Figure 5.9: Schematic of microwave heating. .......................................................................... 78

Figure 5.10: Ignition time of a 2Al-Fe2O3-3(Al-Ni) composite with a 0.05 mm embedded

copper wire vs. the length of the copper wire. Copper wires with no composites

sparked at times comparable to the ignition times. ................................................ 83

Figure 5.11: Spark time vs length for aluminum wire in air. ...................................................... 84

Figure 5.12: Spark time vs length for copper wire with diameter of 0.05 mm in argon

atmosphere. ............................................................................................................. 85

Figure 6.1: Consolidating of nanoheater with two copper strips embedded on a copper sheet.88

Figure 6.2: Mono-disperse SN100C® droplets produced by the UDS process......................... 89

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Figure 6.3: Schematic of joining setup of a copper sheet to a solder layer. ............................. 89

Figure 6.4: Cross section of a reactive composite consolidated with two embedded copper

wires on top of a copper sheet. ............................................................................... 90

Figure 6.5: Cross section of soldered copper using 2Al-3CuO-1(Al-Ni) reactive composite. . 93

Figure 6.6: Ultrasonic abrasive activation of aluminum surface. ............................................. 94

Figure 6.7: Sedimentation sprinkling of corundum particles on solder discs. .......................... 96

Figure 6.8: Schematic of fabrication of an aluminum-solder-aluminum tensile specimen (a)

ultrasonic surface activation, (b) solder reflow. ..................................................... 97

Figure 6.9: Solder reflow thermal profile. ................................................................................ 98

Figure 6.10: Schematic of tensile shear test. ............................................................................... 99

Figure 6.11: Cross sections of etched solder layer for (a) as consolidated, (b) after reflow. ... 101

xvi

Figure 6.12: Cross section of soldered aluminum activated with 0.3 μm alumina at (a) 1 g/m2,

(b) 2 g/m2, (c) 3 g/m

2, (d) 4 g/m

2, and (f) 1g/m

2. ................................................. 103

Figure 6.13: SEM micrographs of the cross section of aluminum-solder specimens activated

with 0.3 μm alumina: (a) gaps at interface (x = 1 g/m2), (b) well bonded interface

(x = 2 g/m2), (c) alumina colonies remaining at interface (x = 4 g/m

2). ............... 104

Figure 6.14: Cross section of a control aluminum-solder specimen reflowed without surface

activation. ............................................................................................................. 105

Figure 6.15: Cross sections of soldered aluminum activated with 1 μm alumina at (a) 1 g/m2 and

(b) 2 g/m2. ............................................................................................................. 106

Figure 6.16: Fraction of the interface joined of aluminum-solder-aluminum specimens vs.

concentration of alumina used for surface activation. .......................................... 107

Figure 6.17: Stress - displacement curve of an aluminum-solder-aluminum specimen activated

with 0.3 μm alumina at 2 g/m2. ............................................................................ 108

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Figure 6.18: Normalized Joint strength of aluminum-solder-aluminum specimens vs.

concentration of 0.3 μm alumina used for surface activation. ............................. 109

Figure 6.19: Cross section of aluminum sheets solder-joined with ultrasonic abrasive surface

activation using 0.3 μm alumina at x = 1 g/m2. .................................................... 110

Figure 6.20: A specimen activated with 0.3 μm alumina at x = 2 g/m2 shows fracture path

through the solder layer. ....................................................................................... 111

Figure 6.21: Schematic of joining setup of a copper sheet to a solder layer. ........................... 112

Figure 6.22: Cross section of soldered aluminum activated with 0.3 μm alumina at 2 g/m2. .. 113

xviii

List of Tables

Table 3.1: Reactive composites composition and their heat outputs for the reactions indicated

[82, 83] ..................................................................................................................... 15

Table 3.2: Four different methods used for mixing .................................................................. 23

Table 4.1: Parameters used for early-stage simulations............................................................ 60

Table 5.1: Voltages at different levels of thermocouple welder. .............................................. 67

Table 5.2: Resistance measurement results of nanoheaters of various compositions and mixing

procedures across their diameters using first method. ............................................. 73

Table 5.3: Resistance and gap distance measurement results ................................................... 76

Table 5.4: Resistance of 2Al-3CuO-4(Al-Ni) reactive composites with embedded two copper

wires and voltages .................................................................................................... 77

1

1. Introduction

Temperature is a major variable for the thermodynamic state of materials, as seen in

dG=-SdT+VdP, therefore, controlling the temperature of materials has been the most critical

requirement in materials processing and manufacturing [1]. Various methods have been

developed and used for the application of heat in materials processing and manufacturing. These

conventional heating methods are largely categorized into four types: (1) fuel-based heating, (2)

electricity-based heating, (3) steam-based heating, and (4) hybrid process heating technologies

[2], most of these conventional heating methods, however, are developed for macro-scale heating

of bulk material, and as such do not apply to the rising need for local heating of small volumes of

materials and micro-scale parts. The ability to provide controlled heating to small designated

volumes or parts of materials is also critical to innovating manufacturing for drastically reduced

energy consumption and impact to the environment and hence for the ultimate sustainability of

human activity.

Efforts are being made to meet the demand for methods of controlled local heating of

micro-scale objects. Among such heating methods, use of self-heating materials as heat sources

for microjoining and local heat treatment is at the center of attention [3, 4]. Self-heating

2

materials are characterized by the exothermic reaction among their constituents. The constituents

can be combinations of metals, metal oxides, and sometimes polymers, that generate large

reaction enthalpy when reacting. Among the most widely studied combinations are bimetallic

Al-Ni and Al-oxide thermite combinations.

Reactive composites suitable for local heating must necessarily be small-sized, and be

composed of even smaller-sized constituents. Ultimately, ideal reactive composites should be

comprised of nano-sized particles or layers of reactive constituents. Such nano-structural reactive

composites, referred to as nanoheaters [5], has been fabricated by PVD [6, 7] and CVD [7]

which are used mainly for fabricating thin film of alternating nanolayers of reactive metallic

constitutes and more recently by ultrasonic powder consolidation (UPC) [8-11] an extension of

ultrasonic welding [12], which applies to powders of various types of constituents. While PVD

and CVD methods are much more studied for the fabrication of nanolayered reactive composites,

they suffer from high equipment cost in cost-conscious manufacturing. UPC, on the other hand,

may not apply to the production of ultrafine-sized nanoheaters, but has advantage in both cost

and versatility [8]. Thus, for UPC has been studied mainly for the fabrication of bimetallic Al-Ni

composites, although limited reports on the fabrication of thermite nanoheaters have been made

[3, 13].

In the present investigation, fabrication of a new type of nanoheaters based on hybrid

compactions of bimetallic Al-Ni and thermite Al-oxide reactive composites is motivated by the

3

opportunity to combine the excellent ignitability of bimetallic Al-Ni composites and the large

reaction enthalpy of thermite in the reactive composite. The main objective of this research is to

study the processing-structure-property relationship of hybrid nanoheater structures and the

application of these novel structures to microscale joining. The specific objectives of this

research are to (1) establish optimum processing conditions for the fabrication of hybrid

bimetallic-thermite nanoheaters by UPC, (2) study the ignitability of fabricated reactive

composites and (3) study the functionality and reliability of the joining/interconnects.

The fabricated hybrid reactive composites were characterized for their microstructures

and ignition in both continuous heating and spot ignition methods. The hybrid composites were

tested as heat sources for microjoining of aluminum/copper sheets with solder. To this end an

effective flux-less method was realized for aluminum soldering in a novel process that combines

ultrasonic abrasive surface activation of aluminum at room temperature and subsequent heating

for reflow. Use of reactive composites allows for exact content of the amount of heat needed for

reflow, and hence produce good solder joints.

4

2. Background

This chapter reviews previous studies on the application of reactive metal matrix

composites, methods of reactive composite fabrication from powder constituents, techniques for

igniting reactive composites and methods for the joining of aluminum.

2.1 Reactive Composites and Their Application in Joining

Nanoheaters are defined as reactive composites made of materials, nano-sized at least in

one dimension that output heat in calculated amounts through exothermic reactions of the

constituents [3, 5]. Reactive composites are largely categorized into the bimetallic (e.g., Al - Ni),

thermite (e.g., Al-metal oxide) and metal-polymer (e.g., Al - TeflonTM

) types [14]. They have

been used as heaters, propellants [15-17] and explosives, as well as for self-propagating high-

temperature synthesis (SHS) [18-25]. Applications of reactive materials began in the welding of

rail sections where Al-iron oxide thermites were traditionally used [26, 27] and have recently

extended to microjoining [28, 29] and soldering in microelectronics [30] as nanostructured

reactive materials suitable for localized heating became available. For example, pressurized

5

mixtures of Ti-2B and Ti-C between two Mo sheets have been used as heat sources for ceramic-

to-metal joining [31]. Au-coated stainless steel sheets brazed with self-heating nanoscale

multilayer Al/Ni foils have exhibited shear strengths as high as 48 MPa, which substantially

exceeds that of conventional solder joints which is typically about 38 MPa [28, 32]. More

recently, exothermic blends of 14Al-3CuO-Ni were used for joining 1100 aluminum alloy sheets

[33].

Due to the metastable nature of reactive composite materials, their fabrication requires a

method that does not employ high temperature processing. Techniques investigated for preparing

powder-based nanoheaters include cold isostatic pressing and sintering [34], high-pressure shock

activation [35], arrested reactive milling [36], sol gel synthesizing [37], and more recently cold

spray (CS) [38] and ultrasonic powder consolidation (UPC) [5]. UPC is a new additive

manufacturing technique in which powders, instead of sheets or wires, are metallurgically

consolidated by the action of ultrasonic vibration at low to moderate temperatures, typically

within a second [8], and as such provides an enabling technique for the consolidation of reactive

powder mixes, as reported for Al-Ni bimetallic and Al-Fe2O3 and Al-Cu2O thermite composites

[39].

In the present work, different types of reactive composites were fabricated by UPC,

including a new type of hybrid bimetallic-thermite nanoheaters (2Al-Fe2O3-x(Al-Ni) and 2Al-

3CuO-x(Al-Ni) and reference composites of compositions 2Al-Al2O3-x(Al-Ni)). The hybrid

6

compositions were chosen as they provide an opportunity to combine the excellent ignitability of

bimetallic Al-Ni nanocomposites and the large heat output of thermites in one reactive

composite. The microstructure and ignitability during continuous-heating of fabricated reactive

composites were characterized, and their applicability to microjoining was investigated with

special interest in enabling soldering for aluminum parts. The 2Al-Al2O3-x(Al-Ni) composites

were added to investigate the ignition characteristics of composites with and without a thermite

reaction.

2.2 Ignition Method

There are many different methods for initiating the reactions in energetic materials, such

as by applying hot spots [40, 41], laser pulses [42, 43], microwaves [44], and plasma [45].

Conventionally, reactive materials are ignited either in a furnace or by the application of a hot

spot [46] or laser [47].

Spark plasma sintering (SPS) [48-50] is a rapid sintering technique, in which a pulsed DC

current passes through the powder compact to effect the consolidation of the powder. [48] These

method is used for the fabrication of micro- and nano- ceramics [50-52]. For conductive

materials joule heating also is very effective in initiating SPS.

7

Another method used for starting SPS is to subject the reactive powder compact to

microwaves. This is typically used for ceramic preparation [52], since most ceramics do not

absorb microwave at room temperature. However their absorption of microwaves increases at

elevated temperatures [53], which in turn increases the temperature of materials. The power

absorbed per unit volume is expressed by [54]:

𝑃 = 𝜎|𝐸|2 = 2𝜋𝑓𝜀0𝜀ef𝑓" |𝐸|2 (1.1)

where E is the electric field magnitude, σ is the conductivity, f is the microwave frequency, 𝜀0 is

the permittivity of free space, 𝜀𝑒f𝑓" is the relative dielectric factor.

The initiation of ceramic sintering by microwave heating is based on hot spot formation

in the specimens [55]. Microwave heating is also used for the sintering of metals to form

intermetallic components [56]. Hot spots form in a material when two waves match up, so that

the distance between two hot spots would be half of the wavelength. The wave length of

conventional commercial microwave ovens with a frequency of 2.45 GHz is 12.24 mm, so

microwave sintering cannot be effective on specimens smaller than half of the wavelength.

Recently another method is used to ignite energetic materials [57], in which a thin wire of

a metal with high resistivity, such as platinum [58], was coated with a small amount of energetic

materials and then a voltage pulse is applied to the wire to cause Joule heating. Impact ignition

[59] is another method for igniting nano- and micro-energetic materials. An impact ignition setup

8

contains a drop mass which is released from different heights and makes contact with a loose pile

of energetic material.

In this study a spark method and continuous heating in a furnace were used to study the

ignitability and ignition characteristics of reactive composites. Moreover, two new methods,

based on Joule heating and microwave ignition, were introduced to provide effective means for

igniting nanoheaters in industrial applications such as micro-scale joining.

2.3 Aluminum Soldering

Aluminum and its alloys, being light materials with high corrosion resistance and thermal

and electrical conductivity, are used in many sectors of manufacturing. Increased utilization of

aluminum is called for particularly in automobile industry where strict fuel economy standards

mandate drastic reductions of vehicle weight. The automobile components that have used

aluminum alloys include roofs, doors, hoods, frames, wheels and engine blocks [60].

A key to further increasing the aluminum utilization in manufacturing lies in the

availability of economical and dependable means to join aluminum to other materials. The

joining technologies for aluminum are categorized largely into (1) fusion welding such as gas

metal arc welding (GMAW), tungsten inert gas (TIG) welding, plasma welding, electron beam

9

welding, laser welding and resistance spot welding (RSW), (2) solid-state joining such as

diffusion bonding, ultrasonic welding and friction stir welding (FSW), (3) brazing, (4) soldering

and (5) non-metallurgical joining such as fastening and adhesive joining [60, 61].

Joining of aluminum, however, is intrinsically difficult due to the properties unique to

aluminum, particularly the unavoidable oxide formation on aluminum surface in air before and

during joining [61, 62]. In fusion welding, prior removal of surface oxide and contaminants is a

must. Aluminum fusion welding also suffers from the high costs of shielding gas, e.g., high-

purity helium and argon gases, crater cracking caused by the high thermal expansivity and

burnthroughs at low welding speeds due to aluminum’s high thermal conductivity [63].

Formation of brittle intermetallics adds to the problems in welding aluminum to other materials

such as steel [64-66]. RSW, a fusion welding method widely adapted in automobile body

assembling, also requires surface oxide control and is largely limited to sheet joining [67].

Diffusion bonding of aluminum requires both surface oxide removal and atmosphere control at

the joining temperature [68]. Ultrasonic joining and FSW are less sensitive to prior surface

conditions, but the former applies only to small parts [69], while the latter must employ

expensive product-specific tools and equipment [70]. Brazing is widely used to join non-heat

sensitive aluminum parts [71] but still requires thorough prior surface cleaning as well as fluxing

to remove/convert oxide at the brazing temperature, often just below the melting point of

aluminum. Soldering provides a means to join aluminum parts without much thermal

degradation or distortion of parts, but normally requires pre-treatment such as pre-plating [72],

oxide removal with flux [73] or other media, e.g., fluorine containing plasma [74], which adds to

10

the cost and also impacts the environment. The use of flux may leave residue in the joint that

causes corrosion as well.

Soldering, nonetheless, can potentially expand its application to wider ranges of

aluminum parts and products if it can be done flux-lessly and without extensive pre-treatment.

Flux-less aluminum soldering has been investigated via various methods [73] including

mechanical rubbing of aluminum surface with molten solder [75], ultrasonic bath soldering [76-

78], and thermal spraying [79]. However, these reported flux-less soldering processes employ

zinc-based solders to assure good wetting in reflow over the aluminum surface [73, 80]. This

gives rise to problems such as sagging, warping, softening, re-alloying and hot cracking of the

base metal. Moreover, zinc solders are not applicable to heat-sensitive parts such as electronic

chips and age-hardened aluminum parts [81]. Use of tin-base solders reduces the reflow

temperature but few reports are found on the flux-less soldering of aluminum with tin-base soft

solders.

An economical and effective method was developed in this work for flux-less direct

soldering of aluminum with tin-base solders in which fluxing is replaced with in-process

ultrasonic abrasion to activate the aluminum surface for fresh metal-to-metal contact with the

solder is presented in this work.

11

3. Ultrasonic Consolidation of Reactive

Composites

As explained in Chapter 2, the reactive composite fabrication requires a method that does

not employ high temperatures. Ultrasonic powder consolidation (UPC) is a new additive

manufacturing technique in which powders, instead of sheets or wires, are metallurgically

consolidated by the action of ultrasonic vibration at low to moderate temperatures, typically

within a second [8], and as such provides an enabling technique for the consolidation of reactive

powder mixes. In this study a new type of reactive composites with hybrid bimetallic-thermite

Al-Ni-Fe2O3 and Al-Ni-CuO compositions were introduced. The hybrid compositions were

designed to take advantage of the large heat output of the Al-metal oxide thermite reaction and

the low ignition temperature of the Al-Ni exothermic reaction in one reactive composite. UPC

was used fabricate specimens of reactive hybrid bimetallic-thermite composites from nanothick

Al and Ni flakes and Fe2O3, CuO and Al2O3 nanoparticles. This chapter describes the

consolidation process of the reactive composites and their thermal behavior. This work is also

presented in the publication by the author: S. Gheybi-Hashmebad and T. Ando, “Ignition

Characteristics of Hybrid Al-Ni-Fe2O3 and Al-Ni-CuO Reactive Composites Fabricated by

Ultrasonic Powder,” Combustion and Flame, 2015, Volume 162, Issue 4, p. 1144.

12

3.1 Experimental Procedure

3.1.1 Materials

The materials used in this study are (1) Al and Ni nanoflakes with sub-micron thicknesses

(100-300 nm), shown in Figure 3.1 (2) thermal plasma synthesized Fe2O3 and CuO

nanopowders, 20-40 nm and 30-50 nm, respectively, shown in Figure 3.2. The Al and Ni

nanoflakes were produced by a proprietary ball milling technique at Fukuda metal foil and

powder Co. Ltd. To provide control specimens that contained a dummy oxide, fine Al2O3

particles (50 nm), were obtained from Alfa Aesar, Figure 3.2(c).

The as-received Fe2O3 and CuO nanoparticles were agglomerated, whereas the Al and Ni

flakes and the Al2O3 particles were loose. The Al and Ni flakes came coated with stearic acid to

prevent pyrophoric reactions. Therefore they were first rinsed in agitated ethanol for about 2

minutes and dried in air. The Al flakes were then blended with Ni flakes, or with Ni flakes and

metal oxide particles (Fe2O3, CuO or Al2O3) in the molar ratios presented in Table 3.1.

13

Figure 3.1: SEM pictures of as-received flakes (a) Ni and (b) Al.

(a)

(b)

14

Figure 3.2: SEM pictures of as-received oxide porticles: (a) Fe2O3 nanoparticles, (b) CuO

nanoparticles and (c) Al2O3 particles.

15

Table 3.1: Reactive composites composition and their heat outputs for the reactions

indicated [82, 83]

Heat output

Reaction

Per reaction

(kJ)

Per unit

composite

volume (kJ/m3)

Bimetallic Al+Ni → AlNi 120 0.9×1 07

Thermite 2Al+Fe2O3 → Al2O3+2Fe 850 2.2×1 0

7

2Al+3CuO→ Al2O3+3Cu 1210 2.7×1 07

Hybrid bimetallic

thermite

2Al+Fe2O3+1(Al+Ni) → Al2O3+2Fe+1AlNi 970 2.0 ×1 07

2Al+Fe2O3+2(Al+Ni) → Al2O3+2Fe+2AlNi 1110 1.8 × 107



2Al+3CuO+1(Al+Ni) →Al2O3+3Cu+1AlNi 1340 1.8 × 107




2Al+Al2O3+1(Al+Ni) →2Al+Al2O3+1AlNi 120 2.1 × 106




16

3.1.2 Experimental Setup

Fabrication: A STAPLA Condor® ultrasonic welding unit is used for the UPC experiments. The

schematic of the UPC experimental setup is presented in Figure 3.3. It operates at a maximum

power of 3 kW and a fixed frequency of 20 kHz, which was used in this work. The amplitude of

vibration is adjustable by changing the level of power but it was fixed at 9 μm in this study. The

UPC setup is connected to a controller, which transmits high frequency signals to the converter

that creates ultrasonic vibration. The setup also contains a heater plate which allows for

performing UPC experiments at elevated temperatures.

Figure 3.3: Schematic view of an ultrasonic welder modified for using consolidates powder

by die and punch.

17

The sonotrode is made of high speed tool steel AISI T9, thus the maximum UPC

temperature is limited below about 833 K in order to avoid exceeding the tempering temperature

of the high speed tool steel. The sonotrode tip is in the size of 4 × 4 mm and has a 14 × 14 grid of

knurls with the dimensions as shown in Figure 3.4.

Figure 3.4: Schematic diagrams showing the sonotrode tip.

Two cartridge heaters (TUTCO, 9.5 mm diameter, 51mm length, 400 W) and a K-type

thermocouple (OMEGA Model SP-GP-K-6) probe are inserted in a stainless steel plate, as seen

in Figure 3.4. 50 mm thick alumina-silica fiber sheets are used for insulating the heater plate.

18

Figure 3.5: A schematic view of the heater plate.

The heater plate holds a K-type thermocouple that is connected to a PC through a control

box and DAQ board. The DAQ system is used for obtaining the temperature data with a K-type

thermocouple inserted in the heater plate. The heaters and the thermocouple are connected to the

control box (Figure 3.6) which is controlled by a computer equipped with a National Instruments

(NI) PCI-6035E multifunction data acquisition (DAQ) card through a NI BNC-2110 connector

block. The control box consists of a solid state relay (SSR), a fuse block with a 250 V, 5 A, and

an Omega FHS-7 finned heat sink. A wiring diagram is given in Figure 3.7. A computer

program, coded on LabView 8.6, obtains the temperature data from the thermocouple and sends

signals to the heaters to heat the plate to the set temperature. This program is set to record the

temperature data with a sampling rate of 1000 data points per second. Additional K-type

thermocouples are used to monitor the specimen temperature during experiments.

Cartridge

heaters

Thermocou

ple probe

19

Figure 3.6: The control box for the heater plate.

Figure 3.7: Wiring diagram for the heater control box.

Vac

Temperature

Controller

dc input

SSR

Control

Side

0 to 5 Vdc,

Typically

4

3

1

2

Load

Side

Heater

Vac

Fast Blow

Fuse

Fuse

On/Off

button

Solid-state

relay

Heat sink

20

The UPC experiments employed a die - punch arrangement, also depicted in Figure 3.3.

The die and the punch were made of 2 mm and 3.5 mm thick stainless steel sheets, respectively,

both of which had a matching diameter of 4.1 mm. The die was first placed on the heater plate

kept at the desired consolidation temperature. Then, about 0.1 g of powder mixture was placed in

the die and held under a normal pressure of 100 MPa applied through the sonotrode and the

punch. This heated the powder mixture to the set consolidation temperature typically in about 60

s. As soon as the powder compact reached the consolidation temperature, 20 kHz in-plane

ultrasonic vibration was applied to the powder compact for 1 s at an amplitude of 9 μm through

the sonotrode and the punch while keeping the compact under a normal pressure of 100 MPa.

The consolidated specimen was removed from the die immediately after the vibration was turned

off.

Metallographic characterization: The consolidated specimens were mounted in epoxy, and

ground on SiC abrasive papers to prepare them for metallographic observation. Several coarse to

fine grit sizes (240, 320, 400, 600, 800, 1200, 1500, 2000) were used. Following the grinding,

the specimens were polished on rotating wheels with fine alumina powder of sizes 1, 0.3 and

0.05 μm.

The mounted specimens were metallurgically characterized under an Olympus VANOX-

T optical microscope with magnification up to 1000x. A JEOL JSM-6360 scanning electron

21

microscope (SEM) (3 nm resolution at 30 kV accelerating voltage) was used for higher

magnifications.

3.1.3 Continuous ignition test

Consolidated nanoheaters were subjected to continuous heating tests to investigate their

ignition characteristics, with particular interest in determining the exact sequence of events that

lead to the ignition of hybrid bimetallic thermite reactive composites during continuous heating.

A National Instruments PCI-6035E data acquisition (DAQ) card with a rate of 1000 data points

per second, through a K-type thermocouple touches the specimen, was used for recording the

temperature profile of the reactive composites. The continuous heating ignition test setup is

schematically illustrated in Figure 3.8. The system consists of a heater plate connected to a PC

through the DAQ board. The heater is placed under enclosed Ar bath. The DAQ system obtains

temperature data with a K-type thermocouple inserted to the heater and another K-type

thermocouple attached to the specimen. The heater plate is set to heat up to 933 K, which is the

melting temperature of aluminum, at a heating rate of 125 K/minute. The PC records the changes

in heater and specimen temperatures and T-t diagrams are created.

22

Figure 3.8: Schematic representation of the continuous heating ignition test setup.

3.2 Results and Discussion

3.2.1 Mixing Process

To achieve good mixing in the powder blends, different combinations of dry mixing in

air, rotary mixing in ethanol and sonication in ethanol were tested. These methods are

23

summarized in Table 3.2. The last two methods (III and IV) use the same mixing time, but differ

in the sequence of mixing.

Table 3.2: Four different methods used for mixing

Method Dry mixing without

any mixing medium

Rotary

Mixing in

ethanol

Sonication in

ethanol

Rotary

Mixing in

ethanol

Sonication in

ethanol

Rotary

mixing in

ethanol

I 60 min 120 min - - - -

II 60 min 60 min 150 min - - -

III 60 min 120 min 300 min 20 min - -

IV 60 min 60 min 150 min 60 min 150 min 20 min

Figure 3.9 shows the cross sections of specimens ultrasonically consolidated with the

different mixing processes at 573 K under 100 MPa pressure for 1s. Figure 3.9 (a) shows the

specimen fabricated using the blend prepared by Method I in which Ni particles are still

agglomerated. To improve the mixing process, additional 300 minutes of sonication in ethanol

was added to the mixing process. Figure 3.9 (b) shows the cross section of a specimen

consolidated by Method II. It shows that sonication in ethanol has a significant effect on

deagglomeration of nanoparticles and nano-flakes. In Method III rotary-mixing and sonication

in ethanol are increased to 120 minutes and 300 minutes, respectively. Then it is followed by 20

minutes of rotary mixing in ethanol before drying. Figure 3.9 (c) shows the specimen that was

consolidated using the blend prepared by Method III. It shows better deagglomeration of Ni

nano-flakes and nanoparticles clusters. Figure 3.9 (d) shows the cross section of a specimen that

24

was consolidated using the mixture prepared with Method IV. It is obvious from the picture that

a better distribution is achieved by this method. Compering Figure 3.9(c) and (d) shows changing

the sequence of mixing steps improves the distribution of all constituents of particles. The rotary

mixing between sonication steps in Method IV, mixed the particles more homogeneously.

Figure 3.9: Effect of mixing method on distribution of particles, particles mixed with (a)

method I, (b) method II, (c) method III and (d) method IV.

With the results of these systematic experiments, the blended powders were dry-mixed in

a cylindrical container 40 mm in diameter and 38 mm in depth rotating horizontally at 750 rpm

(a) (b)

(c) (d)

25

for 60 minutes with no mixing aid medium. Then, the particles were further rotary-mixed for an

additional hour in ethanol. This was followed by ultrasonic sonication in ethanol for 2.5 hours in

order to reduce the agglomeration of nanoparticles, and finally by 60 minutes of additional

rotary-mixing and 300 minutes sonication, before drying in air and final pulverizing and dry-

mixing for 20 minutes. This procedure was used for the UPC experiment of all the hybrid

bimetallic thermite composites.

3.2.2 Microstructure

Figure 3.10 shows the microstructure of Al-Ni bimetallic and Al-CuO thermite

composites, both consolidated for 1 s at 523 K under 100 MPa, which were fabricated as binary

references, as reported in previous study [84]. Aluminum, the softer phase, took most of the

plastic deformation needed for the densification of the powder compact. The bimetallic

composite exhibits a good distribution of Ni flakes (light) in the Al matrix (dark), Figure 3.10(a),

although some of Ni flakes are still present. A similar degree of stacking is considered to remain

with Al flakes as well. Nonetheless, in overall distribution of Ni flakes is fairly uniform. The Al-

CuO thermite composite, Figure 3.10(b), also exhibits a uniform distribution of CuO particles in

the Al matrix, but with same degree of agglomeration that leaves small colonies of CuO along

same of the interface between Al flakes. Both composites exhibit no porosity or unfilled voids,

26

suggesting that the UPC conditions applied were sufficient for full densification in both

composites.

Figure 3.10: SEM image of an (a) Al-Ni (b) Al-CuO reactive composite consolidated for 1

s at 523 K under 100 MPa uniaxial pressure.

27

Figure 3.11 shows SEM images of consolidated specimens of hybrid bimetallic thermite

2Al-Fe2O3-3(Al-Ni), 2Al-3CuO-1(Al-Ni) and 2Al-Al2O3-3(Al-Ni) composites in their cross

sections. The three composites all exhibit a microstructure consisting of Ni flakes (light) and

oxide particles (dark) distributed evenly in a fully densified and metallurgically bonded Al

matrix (gray) [85]. The oxide particles are located both at the interface of Al and Ni flakes and

within the Al matrix. Despite the long procedure of premixing and sonication used, the prior

agglomeration of Fe2O3 and CuO nanoparticles, Figure 3.11 (a) and (b), somewhat persisted in

the consolidated hybrid composites, forming elongated clusters of Fe2O3 and CuO nanoparticles

along the Al-Ni interface (arrows). The Al2O3 particles in the 2Al-Al2O3-x(Al-Ni) composite

were not as much agglomerated since they were loose initially. The composites with the other

compositions shown in Table 3.1 all exhibited a well-consolidated hybrid composite

microstructure similar to the ones in Figure 3.11.

28

Figure 3.11: SEM images of (a) 2Al-Fe2O3-3(Al-Ni), (b) 2Al-3CuO-1(Al-Ni) and (c) 2Al-

Al2O3-3(Al-Ni) reactive composites, consolidated by subjecting premixed

compacts to ultrasonic vibration for 1 s at 573 K under 100 MPa uniaxial

pressure.

29

3.2.3 Ignition Test

Continuous-heating ignition tests were repeated 10 times for each of the hybrid

composites using the procedure described in Section 3.1.3.

Reaction products: All of the consolidated composites ignited when subjected to continuous

heating at a rate of 2.08 K/s (125 K/min) in Ar atmosphere. The final reaction products in the

2Al-Fe2O3-x(Al-Ni) specimens were identified by XRD as Al2O3, Fe and AlNi, and Al2O3, AlNi

and Cu in the 2Al-3CuO-x(Al-Ni) specimens, Figure 3.12. No XRD patterns are available from

the ignited 2Al-Al2O3-x(Al-Ni) specimens from this work, but previous study on Al-Ni

nanoflake composites [86] suggests that their reaction products would consist of Al3Ni, Al3Ni2,

AlNi, and AlNi3.

Sequence of events leading to ignition: Figure 3.13 shows the change in specimen temperature

Tsp just before and at the time of the ignition of a 2Al-3CuO-3(Al-Ni) composite. We note that

Tsp increased linearly up to 855 K as programmed for the heater plate. Thus, below 855 K

exothermic reactions, such as 3Al+Ni → Al3Ni and Al3Ni+Ni → Al3Ni2 [85], did not cause the

specimen to heat up faster than the heater plate although they might take place at the Al-Ni

interface even below 855 K [84, 85]. Figure 3.14 shows evidence of such solid-state Al-Ni

compound formation in a 2Al-Al2O3-3(Al-Ni) specimen that was heated to 773 K.

30

Figure 3.12: XRD patterns of ignited hybrid composites. (a) 2Al-Fe2O3-3(Al-Ni) and (b)

2Al-3CuO-3(Al-Ni)

31

Figure 3.13: The change in specimen temperature with time of a 2Al-3CuO-3(Al-Ni)

composite consolidated at 573 K for 1 s under 100 MPa (a) acceleration

stage above 855 K followed by melting of aluminum and ignition at 895 K,

(b) Tsp vs. t near/at ignition.

32

Figure 3.14: SEM image of a 2Al-Al2O3-3(Al-Ni) composite continuously heated and held

at 773 K for 5 minutes. Al3Ni formed between Ni flakes and Al matrix.

Above 855 K, the exothermic Al-Ni reactions progressively accelerated, causing the

specimen to self-heat at a higher rate of about 10 - 30 K/s. The exothermic Al-Ni reactions in the

acceleration stage are not likely to involve liquid formation, which, if it happened, would trigger

ignition. We may term the specimen temperature at which the acceleration stage started (855 K

for this composite) the acceleration temperature Tac . Above Tac , the reaction would accelerate

throughout the specimen volume, so that the temperature monitored at the top surface of the thin

specimen should practically represent the specimen temperature. This composite ignited at about

1.3 s into the acceleration stage when Tsp reached 895 K. Thus, we define the ignition

temperature Tig as the specimen temperature reached at the end of the acceleration stage. The Tig

33

so defined is of practical importance as it is the temperature above which a reactive composite

ignites in continuous heating, while the acceleration temperature Tac represents the maximum

temperature at which a reactive composite can remain largely unreacted.

We note that the 2Al-3CuO-3(Al-Ni) composite ignited when Tsp was still well below the

melting temperature of aluminum (933 K), and even below the eutectic temperature of Al and

Al3Ni (913 K). However, this does not mean that the specimen ignited in the absence of melting.

In fact, we note a sharp negative (endothermic) shift of Tsp just before the ignition, Figure

3.13(a), which is indicative of melting. A close look at the change in Tsp near/at the ignition,

Figure 3.13(b), reveals that the endothermic shift was preceded by a slight dip of Tsp over about

1.7 ms, followed by a rapid increase to 970 K which occurred over just 1 ms. The first shallow

dip reflects eutectic melting which would occur at the Al-A3Ni interface, Figure 3.15(a), where

local temperatures likely have hit the eutectic temperature (913 K) [84]. The eutectic liquid

would provide high diffusivity paths for the Al-Ni exothermic reactions to further accelerate,

causing the nearby aluminum to melt. The molten aluminum may the reach some of the CuO

particles at the Al-Ni interface, Figure 3.15 (b), initiating the Al - CuO thermite reaction locally.

The localized occurrence of the thermite reaction caused Tsp to jump up sharply to 970 K at a

rate of about 75,000 K/s, Figure 3.13(b). This resulted in extensive melting of aluminum, as

evidenced by the subsequent sharp endothermic shift to 765 K. The molten aluminum then

34

reacted extensively with CuO, causing Tsp to rise at a very high rate exceeding 600,000 K/s

(ignition).

Figure 3.15: Schematic illustration of liquid formation (a) at the reaction front and (b) its

effect on the onset of thermite reaction.

35

Figure 3.16(a) shows the Tsp - t plot of a 2Al-Fe2O3-1(Al-Ni) specimen near and at its

ignition. The acceleration and ignition temperatures of this composite were, respectively, 885 K

and 933 K. The acceleration stage lasted for about 1.7 s at a heating rate of 10 - 40 K/s. This

thermograph also exhibits an endothermic shift just before the ignition. The enlarged view in

Figure 3.16(b) shows no distinct spike of Tsp prior to the endothermic shift. Instead, the

specimen temperature had already reached the melting temperature of aluminum and somewhat

fluctuated over a few ms before it dipped to 827 K (endothermic shift). Thus, at the end of the

acceleration stage, the remaining aluminum started melting, but rather slowly at a rate

comparable to those of the exothermic Al-Ni reactions that also continued, keeping Tsp around

the melting temperature of aluminum. The melting of aluminum then accelerated as manifested

by the endothermic shift to 827 K, at which the Al - Fe2O3 thermite reaction was triggered,

causing Tsp to increase sharply at an initial rate of 400,000 K/s. Eutectic melting did not play as

distinct a role in the this specimen as in the 2Al-3CuO-3(Al-Ni) specimen, but might have

contributed to the increase of Tsp in the late acceleration stage.

36

Figure 3.16: The change in specimen temperature with time of a 2Al-Fe2O3-1(Al-Ni)

composite consolidated at 573 K for 1 s under 100 MPa (a) acceleration stage

above 885 K followed by melting of aluminum and ignition at 933 K, (b) Tsp

vs. t near/at ignition.

37

To further investigate the sequence of events near/at the ignition, additional continuous-

heating ignition tests were conducted with composites of 2Al-Al2O3-x(Al-Ni) in which Al2O3, a

dummy oxide, is substituted for the Fe2O3 or CuO in the hybrid bimetallic-thermite composites.

Since Al2O3 does not react with the metals, these specimens are essentially bimetallic Al-Ni

composites, but with varying amounts of alumina particles consolidated with Al and Ni flakes.

Figure 3.17(a) shows the change of Tsp with time of a 2Al-Al2O3-3(Al-Ni) composite. As in

Figure 3.13 and Figure 3.16 the deviation of Tsp from the plate temperature above 785 K reflects

accelerated reactions of 3Al+Ni → Al3Ni and Al3Ni+Ni → Al3Ni2. During the acceleration stage

(which lasted for 5 s), Tsp increased at an average rate of about 10 K/s, which was slightly lower

than the heating rates during the acceleration stages of the 2Al-3CuO-3(Al-Ni) and 2Al-Fe2O3-

3(Al-Ni) composites. Ignition occurred when Tsp reached 829 K, but with no sharp endothermic

shift preceding the ignition.

38

Figure 3.17: The change in specimen temperature with time for 2Al-Al2O3-3(Al-Ni)

composite consolidated at 573 K for 1 s under 100 MPa (a) acceleration

stage above 783 K followed by melting of aluminum and ignition at 829 K,

(b) shallow endothermic dip just before ignition.

39

A closer look at the thermograph, Figure 3.17(b), however, reveals a shallow endothermic

shift over about 0.3 s just before the ignition. Thus, the local (reaction-front) temperatures did

reach a temperature at which liquid was formed while Tsp was still 829 K. This liquid, which

could come from eutectic melting between Al and Al3Ni and melting of aluminum, further

boosted the 3Al+Ni → Al3Ni and Al3Ni+Ni → Al3Ni2 reactions and could also activate

additional exothermic reactions, such as Al3Ni2 + Ni → 3AlNi, Al3Ni + 2 Ni → 3AlNi and Al +

Ni → AlNi [85, 86], causing the specimen temperature to increase more rapidly (ignition),

Figure 3.17(b). The ignition temperature of this 2Al-Al2O3-3(Al-Ni) composite (829 K) was

determined by linear extension of the specimen temperature over the endothermic shift as shown

in Figure 3.17(b). No spike or subsequent drop of Tsp occurred in this composite because the

reactions among Al, Ni and their compounds were much slower than thermite reactions. The

average heating rate just after the ignition was only about 1000 K/s, which is two orders of

magnitude lower than those observed with the hybrid bimetallic-thermite 2Al-3CuO-3(Al-Ni)

and 2Al-Fe2O3-3(Al-Ni) specimens.

Dependence on x: The 2Al-Fe2O3-x(Al-Ni), 2Al-3CuO-x(Al-Ni) and 2Al-Al2O3-x(Al-Ni)

composites all exhibited distinct acceleration and ignition temperatures, both of which decreased

with increasing x, Figure 3.18. Ten experiments were performed for each of the experimental

conditions investigated, yielding tightly repeated values of acceleration and ignition

temperatures, both within ± 5 K about the values in Figure 3.18. The decrease of Tac with

increasing x, Figure 3.18(a), is understood since composites with higher Al and Ni interfacial

40

areas would have lower onset temperatures for their acceleration stages. The decrease of Tig with

increasing x, Figure 3.18(b), is also expected since Tig is defined as the final specimen

temperature of the acceleration stage which begins and progresses at lower temperatures in

composites with higher x. Figure 3.18(b) also shows that, except for the 2Al-Fe2O3-1(Al-Ni) and

2Al-3CuO-1(Al-Ni) composites, ignition occurred before the specimen temperature reached the

melting point of aluminum. In hybrid bimetallc-thermite composites with high x, the thermite

reaction may be activated in two stages as discussed for the 2Al-3CuO-3(Al-Ni) composite in

Figure 3.13(b): an initial localized stage activated by the liquid formation due to Al-Ni reactions

that brings up the specimen temperature to/above the melting temperature of aluminum (the

spike to 970 K) and a global stage triggered by the resultant extensive melting of the remaining

aluminum. However, hybrid composites with low x, e.g., the 2Al-Fe2O3-1(Al-Ni) in Figure 3.16,

are already at the melting point of aluminum at the end of their acceleration stage where

aluminum can melt with no further jump of Tsp to trigger ignition. It follows that the height of

the spike at the end of the acceleration stage should increase with increasing x.

41

Figure 3.18: Dependence of (a) the ignition and (b) acceleration temperatures of 2Al-

Fe2O3-x(Al-Ni), 2Al-3CuO-x(Al-Ni) and 2Al-Al2O3-x(Al-Ni) composites on

x (bimetallic Al-Ni addition)

42

Figure 3.19 shows this effect for 2Al-3CuO-x(Al-Ni) composites. At x = 1, no spike

appeared, which is consistent with the behavior of the 2Al-Fe2O3-1(Al-Ni) shown in Figure

3.13(b), whereas at x = 2 and 3, a spike appeared and its height increased with increasing x. It

follows that Tig , defined as the specimen temperature at the end of the acceleration stage,

matches with the melting point of aluminum at x = 1, but is depressed below the melting point at

a higher x by an amount that scales with the height to the spike. The thermographs of the 2Al-

Al2O3-x(Al-Ni) composites did not exhibit such a spike nor a sharp endothermic shift, Figure

3.20. This is because the Al-Ni reactions are much slower than the thermite reactions, making the

transition from the acceleration stage into ignition a slower process occurring over time as seen

in Figure 3.18 (b).

Figure 3.19: The height of the spike before ignition depended on x, the bimetallic Al-Ni

addition: (a) no spike appeared at x = 1, (b), (c) spike appeared at x = 2 and 3

with its height increasing with increasing x.

43

Figure 3.20: Specimen temperature vs. time of 2Al-Al2O3-x(Al-Ni) composites. No spike

nor sharp endothermic shift preceded ignition, reflecting the lower rates of Al-

Ni reactions. Modified time (t’) is used to place the three thermographs in one

figure.

Effects of oxide clusters: Figure 3.18 also shows that the 2Al-Fe2O3-x(Al-Ni) and 2Al-3CuO-

x(Al-Ni) composites have much higher Tac and Tig than the 2Al-Al2O3-x(Al-Ni) composites.

This can be explained by the degree of agglomeration of the oxides in the composites, rather than

by the chemistry of the oxides per se. As seen in Figure 3.2, the original Fe2O3 and CuO

nanoparticles were both highly agglomerated, whereas the 0.05 μm Al2O3 particles were loose.

Despite the lengthy mixing and sonication procedure used, the Fe2O3 and CuO nanoparticles still

remained agglomerated in the consolidated composites, forming elongated clusters of oxide

nanoparticles between the metal flakes, Figure 3.11(a) and (b). These clusters are considered to

block the lateral growth of Al3Ni along the Al-Ni interface, thereby retarding the acceleration

44

stage in these composites until higher specimen temperatures were reached. Conversely, in the

2Al-Al2O3-x(Al-Ni) composites, the Al2O3 particles were not agglomerated as much, Figure

3.11(c), allowing for more contiguous Al3Ni growth along the Al-Ni interface at lower

temperatures. The correlation between the Tig of the composites with their respective Tac in

Figure 3.21, which were extracted from Figure 3.17, clearly demonstrate the higher effectiveness

of the Fe2O3 and CuO nanoparticles clusters in retarding the acceleration stage than that of the

Al2O3 particles.

Figure 3.21: High effectiveness of Fe2O3 and CuO nanoparticle clusters in retarding the

acceleration stage over that of Al2O3 particles.

45

Trade off between ignitability and energy output: The hybrid bimetallic-thermite reactive

composites all ignited much below the ignition temperatures of Al - Fe2O3 thermites which are

reported to exceed 1233 K [87]. Therefore, the bimetallic addition, x(Al-Ni), to an Al-oxide

thermite is effective in improving the ignitability of thermite-based composites. Figure 3.22(a)

shows the ignition temperature and energy output of 2Al-Fe2O3-x(Al-Ni) against x where the

energy output is calculated for a unit volume of composites assuming completion of the reactions

in Table 3.1, and linear approximation of the composite density using ρcomp = ∑niρi / ∑ni where

n, and ρ are, respectively, the number of moles and the density of constituent i in the powder

mixture. The value of ignition temperature at x = 0 was found in the literature that reports typical

values for 2Al-Fe2O3 thermites [87]. Clearly, small bimetallic additions of Al and Ni flakes to the

thermite drastically decrease the ignition temperature, i.e., increase the ignitability, while keeping

the heat output at high values. Similar effects are contemplated for 2Al-3CuO-x(Al-Ni)

composites as well, Figure 3.22(b). This confirms that the hybrid bimetallic-thermite reactive

composites combine the excellent ignitability of Al-Ni bimetallic composites and the high heat

output of thermites in single reactive composites.

46

Figure 3.22: Ignition temperature and energy output of (a) 2Al-Fe2O3-x(Al-Ni) and (b)

2Al-CuO-x(Al-Ni) against x. The energy output is calculated for a unit

volume of composites assuming completion of the reactions in Table 3.1.

Literature value (1233 K [87]) is used for x = 0 since 2Al-Fe2O3 specimens

consolidated by UPC did not ignite at the maximum temperature reachable

with the experimental setup.

880

900

920

940

960

980

1000

1020

x (Bimetallic addition)

T (

K)

0 1 2 3 40.6

0.8

1

1.2

1.4

1.6

1.8

2x 10

7

|

H|

kJ

Ignition temperature

Enthalpy (absolute value)

(a)

(b)

2Al‒Fe2O3‒x (Al-Ni)

2Al‒3CuO‒x (Al-Ni)

47

4. Early-Stage Reaction Kinetics Modeling

The high interfacial area in nanoheaters may reduce the ignition temperature compared to

that of their counterparts with low interfacial areas. This is because, as shown in the thermal

analysis in Chapter. 3, the rate of heating due to the exothermic reaction, 3Al+NiAl3Ni and

Al3Ni+NiAl3Ni2 taking place in solid state at the interface of Al3Ni depends on the area of

Al-Ni interface per unit volume of the composite. In a composite with a large Al-Ni interfacial

area, the local temperature at the reaction front in the composites may reach the Al-Al3Ni-L

eutectic temperature (913 K) even when the average specimen temperature is still substantially

below the eutectic temperature. When this happens, the eutectic liquid produced at the reaction

front triggers ignition [3].

A model was previously developed by Erdeniz [86] to predict the early-stage solid-state

reaction kinetics of Al and Ni [86]. This model employs the Avrami equation [88], but with a

modification in a form appropriate for the prediction of the overall reaction kinetics under non-

isothermal conditions encountered in the reactive composite ignition during continuous heating.

This model calculates the increase in volume fraction (Δxv) of the product phase that forms via

the exothermic reactions at a temperature (T) in a small time step (Δt). The calculated value of

48

product volume fraction is then used in a thermal balance equation to obtain the temperature

increase (ΔT) due to the formation of the product phase. The increase in volume fraction in the

next step is calculated for the same time increment Δt, but this time for a temperature of T + ΔT.

Successive iteration of this step yields a T-t plot of a nanoheater for a given starting temperature.

In this study the same approach is used to predict the early-stage solid-state reaction

kinetics of hybrid bimetallic thermite composite structures. For adapting this method to such

hybrid reactive composites, the areas of the Al-metal oxide and/or Al-Ni interfaces need to be

known. This was done by applying an image analysis of the composite microstructures as shown

later in this section.

4.1 Theory

4.1.1 Total Volume of Product Phase

The early-stage solid-state reactions in a hybrid bimetallic thermite nanoheater start at the

intermetallic Al/Ni interface while metal oxide remains unreacted, A (s) + B (s) + C(s) α

(compound) + C(s) + ΔH, Figure 4.1. Assume α disks with a thickness of W, nucleating at a rate

of I, grow laterally along the A/B interface at a growth rate of G. Strictly, the thickness W

49

depends on temperature and time, but further thickening of α disks requires through-thickness

diffusion and as such is sluggish. Thus, it is reasonable to assume that W is determined by the

diffusion-limited growth of α at the triple junction of the α disk and the A-B interface. Thus, it is

a function of the lateral growth rate G and the effective diffusivity D

𝑊 =𝐷

𝐺 (4.1)

where D is the effective diffusivity at the triple point.

Figure 4.1: A schematic showing compound (α) particles forming at the A-B interface.

The volume of the α particles (disks) that nucleate over a time increment between ti and ti

+ dt is expressed by

50

2

/0),(

t

t

iBAii

i

dGWSdtIVttdV (4.2)

where Ii is the nucleation rate at ti, i

BAS / is the area of A/B interface available for α nucleation

per unit volume at ti, V0 is the volume of the A-B-C composite. Integration of Equation 4.2, with

the assumption that α disks impinge only in the lateral growth, i.e. W=D/G, gives the equation

for the total volume of α disks at time t [99].

i

t

t

t

t

i

BAi dtdGTWSIVtV

i

2

/0 )()(

0

(4.3)

by applying the Avrami correction to Equation 4.5 [88], the change in volume fraction, xα, is

calculated by

i

t

t

t

t

j

t

t

t

t

jiBA dtdGTWdtdGIISV

tVtx

i

i

j

i

22

0

/

0

)(.exp)(

)(

0 0

(4.4)

51

Figure 4.2: Compound disks impinging on each other only in lateral growth.

4.1.2 Heat Generation

The solid-state exothermic reaction xA (s) + yB (s) + zC (s) α (compound) + zC (s)

occurring at temperature T, generates heat, ΔH(T), giving by [86]

T

C

P

T

P

T

C

P

B

P

A

P dTzCdTCHdTzCCyCxTHK298K298

K298,

K298

)()(

(4.5)

52

where i

PC ’s are the heat capacities and o

K298,H is the heat of formation for the α compound at

room temperature so the generated heat per unit volume is

V

THTHV

)()(

(4.6)

where Vα is the molar volume of α.

4.1.3 Thermal Balance

In the above model developed at AMPL [86] it is suggested that the thermal balance be

modified for non-adiabatic conditions

dTVCdHdxVTH PLV 00)(

(4.7)

It is assumed that the heat transfers only by natural convection from the top surface. So

the heat loss may be calculated by

0TTdtAhdH SL (4.8)

where T0 is the surrounding temperature or starting temperature, AS is the surface area and h is

convection heat transfer coefficient given by:

53

NuCL

kh

(4.9)

where k is the thermal conductivity of the surrounding, LC is the characteristic length, and Nu is

the Nusselt number:

4/1Ra54.0Nu L (4.10)

and Ra is the Rayleigh number which can be obtained by

Pr)(

Ra2

30

Lv

LTTg C

(4.11)

where β is the coefficient of thermal expansion, g is the gravitational acceleration and defined by

β = 2/(T+T0), Pr is the Prandtl number, and ν is the kinematic viscosity of the surrounding gas

which is calculated by [89]:

𝜈 = 4.7 × 10−11𝑇2 + 8.947 × 𝑇 − 1.734 × 10−5 (4.12)

the Prandtl number is also calculated by [89]:

Pr = −6.05 × 10−11𝑇3 + 2.469 × 10−7𝑇2 − 0.0002611 𝑇 + 0.7798 (4.13)

In this model for the sake of simplicity the temperature of the whole specimen is assumed

to be uniform in each time increment, but in reality hot spots may develop at the reaction fronts

54

i.e., at the triple junctions of A, B and α, as discussed in Section 3. Nonetheless, the model

presented here ignores this effect and assumes uniform heating throughout the specimen volume.

4.1.4 Numerical Computation

Equation (4.4) may be written in the form of summation to carry out step-by-step

computation with a constant time increment Δt.

1

0

22

21

21

0

1330

/ expi

k

i

kj

k

i

kj

k

i

k

i

kj

jkkkBAi GGGItWItSx

(4.12)

The modeling scheme is coded in MATLAB®

and given in the Appendix. First, all the

parameters required to solve Equation (4.12) should be determined. These are 0/ NiAlS , W, G, and

I. The initial Al-Ni interfacial area 0/ NiAlS can be estimated by image analysis as illustrated in the

next section.

Estimation of (S0

Al/Ni) using image analysis: The initial Al/Ni interfacial area ( 0/ NiAlS ) was

determined by using image analysis on a micrograph of actual specimens as in Figure 4.3 which

shows a cross section of a 2Al-Fe2O3-3(Al-Ni) reactive composite.

55

Figure 4.3: Cross section of a 2Al-Fe2O3-3(Al-Ni) reactive composite ultrasonically

consolidated at 573 K under 100 MPa for 1 s.

In order to find the Al-Ni interfacial area, we determined the total length of the Al-Ni

interface per unit area of cross section. First, the Ni particles were detected using ImageJ

software as seen in Figure 4.4(a). Then using the “bwperim” command in MATLAB software,

the boundaries of Ni particles were recognized, Figure 4.4(b). The same technique was used to

determine the Al boundaries in the specimen. Figure 4.5 shows: (a) the Al particles and (b) the

Al particles boundaries in specimen, respectively.

56

Figure 4.4: (a) Ni particles in specimen (b) Ni boundaries in the 2Al-Fe2O3-3(Al-Ni)

specimen.

Figure 4.5: Al particles in specimen (b) Al boundaries in the 2Al-Fe2O3-3(Al-Ni) specimen.

To determine the total length of the Al-Ni interface, the sum of the common pixels in

Figure 4.4(b) and Figure 4.5(b) was first calculated with a simple MATLAB code. The common

pixels were then converted to the total length. In the conversion 6 pixels amounted to 1 m. The

MATLAB code is presented in the Appendix. Figure 4.6 shows the Al-Ni interfacial length

determined on the micrographs of this specimen.

(a) (b)

(a) (b)

57

Figure 4.6: The Al-Ni interfacial length.

The calculated interfacial length was converted to the area per unit volume (m-1

). For this

purpose the specimen was considered as a multilayer structured material. Figure 4.7 shows the

imaginary layers in the specimen. Each layer has a thickness of 1 μm and width of 1 m. The Al-

Ni interfacial length for each layer was calculated using Figure 4.6 and the MATLAB code, in

the Appendix. For a unit volume of the multilayered specimen, each layer has a depth of 1 m, so

the interfacial area for each layer would be equal to the interfacial length per meter. Save was

defined as the average of interfacial area for each layer. The specimen is divided to have 106

layers so the initial Al-Ni interfacial area for the specimen would be

60

/ 10 aveNiAl SS (4.13)

58

Figure 4.7: Considered layers in a specimen.

The lateral growth rate G may be given by an Arrhenius type equation [89]:

Tk

QKG exp (4.14)

where k is the Boltzmann constant and K and Q are the lateral growth factor and the activation

energy, respectively.

59

The thickness of the Al/Ni compound phase may be assumed to be constant as justified in

Section 4.1.1. During these calculations, for the sake of simplicity, W is taken as the half of the

bilayer thickness. Considering N as the nucleation rate and if the nuclei all have the thickness of

W and grow laterally at a rate of G, Equation 4.8 can be rewritten as

2

1

20/ exp1

i

k

kBAi GtNSWx (4.15)

which permits calculating xα as a function of time t. The parameters used for the calculation are

given in Table 4.1.

60

Table 4.1: Parameters used for early-stage simulations.

Parameter Symbol Value Refs.

Initial interfacial area per unit volume for 2Al-Fe2O3-3(Al-Ni) (m-1

) 0/ BAS 3 x 10

6

Compound thickness (m) W 2.5 x 10-7

Number of nuclei per unit area (m-2

) N 1014

- 1016

Time step (s) Δt 10-3

– 10-4

Boltzmann constant (eV K-1

) k 8.617 x 10-5

Lateral growth factor (m s-1

) K 104 [90]

Lateral growth activation energy (eV) Q 1.44 [90]

Molar volume of the compound (m3 mol

-1) Vc 2.4 x 10

-5

Enthalpy of formation at 298 K (J mol-1

) o

K298,H 150624 [82]

Characteristic length (m) LC 10-5

Surface area of the thin film (m2) As 10

-4

4.2 Results and Discussion

Using the parameters given in Table 4.1, the temperature of the specimen was calculated

with different starting temperatures, for both adiabatic and non-adiabatic conditions. It was

assumed that there was no prior reaction in the specimen at the starting temperature and that the

temperature is uniform in the specimen. Figure 4.8 shows the temperature of a 2Al-Fe2O3-3(Al-

61

Ni) composite calculated for the non-adiabatic condition described in Section 4.1.3. The value of

the initial Al-Ni interfacial area per unit volume, 𝑆𝐴𝑙−𝑁𝑖0 , was determined by the image analysis

method described in Section 0. The specimen held at 500 K can reach up to the eutectic point

(913 K) and the liquid phase forms and triggers ignition, in 0.45 s. The ignition temperature is

calculated as the one at which the heating rate exceeds 5 K/ms.

Figure 4.8: Temperature change in a 2Al-Fe2O3-3(Al-Ni) specimen while being held at 500

K, for nonadiabatic condition.

Figure 4.9 shows the results for the adiabatic condition. A specimen held at 550 K can

reach up to the eutectic temperature in just 0.03 s. A specimen held at 500 K can reach the same

point in less than 0.4 s. However, holding the specimen initially at 450 K did not result in a rapid

0 0.1 0.2 0.3 0.4 0.5400

500

600

700

800

900

1000

1100

Te

mp

era

ture

(K

)

Time (s)

550 K

62

temperature increase within 0.5 s. Another run for a longer duration showed that the specimen

can reach the eutectic temperature in 10 s. Similar results were obtained under non-adiabatic

conditions. Figure 4.10 shows the results for three runs starting at 450, 500 and 550 K. Due to

the heat dissipation considered in the calculation, specimens held initially at 500 K and 550 K

are shown to require slightly longer time for ignition to occur than the adiabatic condition. When

started at 450 K, the eutectic temperature is reached in 40 s, a four-fold increase over the time

calculated under adiabatic conditions.

Figure 4.9: Temperature change in a 2Al-Fe2O3-3(Al-Ni) specimen while being held at 450

K, 500 K and 550 K, for adiabatic condition.

0 0.1 0.2 0.3 0.4 0.5400

500

600

700

800

900

1000

1100

1200

Te

mp

era

ture

(K

)

Time (s)

450 K

500 K

550 K

63

Figure 4.10: Temperature change in a 2Al-Fe2O3-3(Al-Ni) specimen while being held at

450 K, 500 K and 550 K, for non-adiabatic condition.

Figure 4.11 shows the specimen temperature simulated for 2Al-Fe2O3-3(Al-Ni) hybrid

nanoheaters starting at 500 K, but with and without consumption of Al-Ni interface by prior

solid-state reactions. The calculated time to ignition is only 0.45 s when the entire interfacial area

is available for Al-Ni reactions at and above 500 K, but it is 0.75 s and 1.43 s when the starting

interfacial area is reduced by 30% and 50%, respectively. Since the time to ignition directly

translates to the acceleration time in continuous-heating ignition, we deduce that prior solid-state

reactions make the acceleration stage longer and might also push it to higher temperatures.

0 0.1 0.2 0.3 0.4 0.5400

500

600

700

800

900

1000

1100T

em

pe

ratu

re (

K)

Time (s)

450 K

500 K

550 K

64

.

Figure 4.11. Temperature change in a 2Al-Fe2O3-3(Al-Ni) specimen held at 500 K while

the effect of solid state reactions below the 500 K was considered.

0 0.5 1 1.5500

600

700

800

900

1000

Time (s)

Te

mp

era

ture

(K

)

S0

Al/Ni0.7S0

Al/Ni0.5S0

Al/Ni

65

5. Ignition Methods for Possible Industrial

Applications

In Chapter 3, a continuous heating ignition test was introduced to study the thermal

behavior of fabricated reactive composites. However continuous heating ignition is not a proper

method industrial applications of reactive composites. In this chapter three different ignition

methods that suit industrial applications are addressed; 1. spark ignition, 2. Joule-heating

ignition, and 3. microwave ignition. The second and third techniques are new approaches for

starting the reaction in nanoheaters which are developed in this study while spark ignition was

used in previous studies [86].

66

5.1 Spark Ignition

5.1.1 Experimental Setup

Figure 5.1 shows a schematic of the setup used for the spark ignition test. An OMEGA

TL-WELD fine-wire welder with an energy output of up to 60 J is employed to provide the

voltage needed to generate a spark. Application of the spark to the nanoheater, placed on a piece

of Ni sheet and held with tweezers, melts down a small portion of the nanoheater and initiates a

reaction which may self-propagate throughout the whole specimen volume if the microstructure

of the specimen is suitable for self propagation [20]. This test was applied to all the reactive

composites of different compositions fabricated in the present study.

The fine-wire welder can operate in three different levels. The voltages and their duration

for each level are presented in Table 5.1. As shown in Table 5.1, the ‘high’ level of the fine-wire

welder provides an electric potential difference of 40 V in 13.6 ms. In this study, all fabricated

reactive composites ignited when subjected to 40 V.

67

Figure 5.1: A fine-wire welder used for spark ignition tests.

Table 5.1: Voltages at different levels of thermocouple welder.

Level Voltage (V) Duration (ms)

Low 38 2.2

Medium 41 6

High 40 13.6

68

5.1.2 Results and Discussions

All reactive composites fabricated in this study were ignitable by spark ignition. Figure

5.2 shows a 2Al-3CuO-1(Al-Ni) specimen subjected to spark ignition. The specimen is seen to

ignite and the self-propagation reaction completed.

Figure 5.2: Different stages of spark ignition of a 2Al-3CuO-3(Al-Ni) specimen

consolidated by subjecting premixed compacts to ultrasonic vibration for 1 s

at 573 K under 100 MPa uniaxial pressure, (a) before applying spark, (b)

application of spark and (c) ignition products.

(a) (b)

(c)

69

The interfacial area has a significant effect on the propagation of the exothermic reaction

throughout the whole volume of the reactive composite. The nano-flakes and nanoparticles to

fabricate reactive composites used in this study (mentioned in Table 3.1) provided a high

interfacial area, so all of the reactive composites transformed completely to the reaction

products. The products of the ignited 2Al-3CuO-x(Al-Ni) specimens were identified by XRD as

Al2O3, AlNi and Cu, as seen for a 2Al-3CuO-1(Al-Ni) specimen in Figure 5.3. The XRD pattern

shows no peak of Al, Ni or CuO, indicating that the 2Al-3CuO-1(Al-Ni) reactive composite fully

reacted by spark ignition.

Figure 5.3: XRD pattern of 2Al-3CuO-1(Al-Ni) ignited by spark ignition method.

70

5.2 Joule Heating Ignition

5.2.1 Experimental Procedure

Another method used for the ignition of the consolidated specimens is Joule heating. As

shown in Figure 5.4(a), in this method nanoheaters with two copper strips were simultaneously

consolidated.

Figure 5.4: (a) Schematic of consolidated specimen with two buried copper stripes. (b) A

possible path of electrons that pass through two copper strips.

(a)

(b)

71

The Al and Ni flakes distributed in a consolidated reactive composite provide a material

with finite resistance between the two copper wires, Figure 5.4(b). The dashed line in Figure

5.4(b) shows a possible path of electrons that pass through the composite when a sufficient

voltage is applied between the two copper strips. Since the resistance of the gap region is much

higher than that of copper strips, heat is generated in the gap region in an amount given by

Joule’s 1st law, i.e., 𝑄 = 𝐼2𝑅𝑡. Two methods were used to measure the resistance of the

nanoheater between the two copper leads. The nanoheaters of 2Al-Fe2O3 and 2Al-Fe2O3-x(Al-

Ni) composites were consolidated at 573 K for 1 s under 100 MPa uniaxial pressure.

First method used for measuring the resistance of nanoheaters: A Fluke 867b graphical

multimeter is used for measuring the resistance of nanoheaters across its diameter. Two lead

wires of the multimeter touch the as-consolidated nanoheater (with no embedded Cu wires) as

shown in Figure 5.5. The resistance of the lead wires was calculated to be 0.24 Ω and was

subtracted from the final measurement.

72

Figure 5.5: Schematic of resistance measurement.

Second method used for measuring the resistance of nanoheaters: Figure 5.6 shows a schematic

of a nanoheater with two embedded copper wires. The resistance is measured for each sample

with a Fluke 867b graphical multimeter. The sample is then mounted and polished to expose the

lead wires. The gap distance between the lead wires is determined via optical microscopy. The

mixing processes used for the specimens described in this section are explained in Section 3.2.1.

73

Figure 5.6: Schematic of resistance measurement method.

5.2.2 Results and Discussion

The resistance values of 5Al-3Ni-Fe2O3 nanoheaters with different mixing conditions and

a 2Al-Fe2O3 reactive composite were measured with the first method (Figure 5.5) across their

diameters. For each reactive composite, ten measurements were made across the circumference

of the reactive composite and the average values are presented in Table 5.2. We see that with

increasing additional mixing steps the resistance decreases. The trend of resistance reduction is

clear in Figure 5.7. We note that the mixing procedure #1 (1 h of dry rotary mixing, 1 h of rotary

mixing in ethanol, 2.5 h of sonication in ethanol) yielded a 5Al-3Ni-Fe2O3 nanoheater with a

high resistance of 109 . However, adding 1 h of rotary mixing in ethanol (procedure #2)

drastically reduced the resistance to 6.5 . Only a slight resistance reduction was attained by

mixing procedure #3 in which 2.5 h of sonication in ethanol is added to procedure #2. The above

74

observations suggest that the sonication in ethanol only helped to deagglomerate the

nanoparticles (as explained in Section 3.2.1) but had little effect on mixing the Al and Ni flakes,

while the additional rotary mixing in ethanol was much more effective in achieving the mixing

of the flakes. The resistance of the Al-Fe2O3 reactive composites was too high to be measured by

the first method.

Table 5.2: Resistance measurement results of nanoheaters of various compositions and

mixing procedures across their diameters using first method.

Nanoheater Mixing procedure

# Details Average

resistance (Ω)

Standard

deviation (Ω)

2Al-Fe2O3-3(Al-Ni)

1

1 hour dry rotary mixing

1 hour rotary mixing in ethanol

2.5 hours sonication in ethanol

108.8 24.8

2



2.5 hours sonication in ethanol


6.5 1.1

3



2.5 hour sonication in ethanol



4.1 1.5

2Al-Fe2O3






- -

75

Figure 5.7: Reduction of resistance by improving the mixing.

The resistance of nanoheaters, with two embedded copper wires (Figure 5.6), was

measured with the second method. The nanoheaters were consolidated using the powders

prepared using mixing procedure #3. The gap between the copper lead wires was then

determined for each nanoheater by optical microscopy. A sample optical microscope image with

the scale is shown in Figure 5.8. The resultant resistance values are presented in Table 5.3 for

2Al-Fe2O3 and 2Al-Fe2O3-3(Al-Ni) composites with their corresponding values of the gap

between the copper wires. We note that the resistance of the 2Al-Fe2O3 composites decreased as

the gap decreased. A similar correlation is found for the 2Al-Fe2O3-3(Al-Ni) composites as well,

although the resistance also depends on the distribution of the Al and Ni nanoflakes which in

1 2 30

20

40

60

80

100

120

Mixing procedure

Re

sis

tan

ce

(

)

Resistance of 2Al-Fe2

O3

-3(Al-Ni)

Trend l ine

76

turn depends on the initial mixing procedure, as well as local distributions of the flakes. We also

note that second method permits measuring the resistance of 2Al-Fe2O3 composites in a

meaningful manner.

Figure 5.8: Example image of the gap measurement method using the optical microscope

image.

Table 5.3: Resistance and gap distance measurement results

Sample Wire gap distance (mm) Resistance

0.2 2.3 (kΩ)

2Al-Fe2O3 1.14 11.5 (kΩ)

1.32 25.5 (kΩ)

0.45 3.3

0.51 7.0

2Al-Fe2O3-3(Al-Ni) 0.6 4.1

1.78 1.7

1.88 9.0

77

2Al-3Cuo-4(Al-Ni) composites with embedded copper leads were ignited by gradually

increasing the voltage applied to the copper leads to determine the voltage at which ignition

occurred. The results are shown in Table 5.4. It is clear that composites with high resistance

require a higher imposed voltage in line with Joule’s 1st law. The value of 𝑉2

𝑅⁄ of the

composites that ignited are all about 1.93 J/s. This verifies that the ignition of the composites was

indeed caused by Joule heating. The composite with low resistance (0.16 Ω) never ignited. Table

5.4 shows that for igniting reactive composites with higher resistance a bigger voltage is needed

(Joule’s first law 𝑄 ∝𝑉2

𝑅), although the reactive composites with very small resistance might

never ignite by this method. They act as a conductive material and will pass the electrons without

producing any noticeable heat.

Table 5.4: Resistance of 2Al-3CuO-4(Al-Ni) reactive composites with embedded two

copper wires and voltages.

M (g) R (Ω) ΔV V2/R (J/s)

0.04 0.16 - -

0.04 33 8 1.93

0.05 206 20 1.94

0.05 852 40 1.87

78

5.3 Microwave Heating Ignition

Another method used for reactive composites ignition employs microwaves. The reactive

composites were consolidated with embedded copper or aluminum wire. Various copper wire

lengths and different atmospheres (air and argon gas) were tried to determine the window for

microwave ignition. A conventional commercial microwave oven of 1000 Watt power was used.

The microwave oven used a 2.45 GHz multimode chamber with a revolving glass plate. In the

ignition test, the sample is placed on the glass revolving plate in the oven.


Reactive composites preparation for microwave experiments: Reactive composite of

compositions 2Al-Fe2O3-3(Al-Ni) which fabricated by UPC with a piece of copper or aluminum

wire embedded in them, Figure 5.9.

Figure 5.9: Schematic of microwave heating.

79

The embedded wire acts as the susceptor of the electromagnetic waves, which induces an

electrical breakdown at an end of the wire. The breakdown occurs due to a corona discharge of

the gas in which the composite with embedded wire is situated [91]. The electric breakdown

potential is given by

𝐸 =𝑄

4𝜋𝜀0𝑟 (5.1)

when 𝜀0 is the permittivity of free space which, for air, is 8.852× 10-12

F/m, Q is the amount of

free charge which could be calculated as a function of the wire volume and r is the radius of the

tip of the end of the wire. In the present work, the wire was cut from continuous wire and as such

had an effective tip radius less than or equal to half of the wire diameter. The electrical

breakdown (E) varies for different atmospheres. It is 3 × 106 V/m for air and 0.6 × 10

6 V/m for

argon.

Electrical breakdown may occur in vacuum by different mechanisms, and is expected to

happen at or near the Schwinger limit [92]. The Schwinger limit is the maximum electric field

above which the electromagnetic field becomes nonlinear. This limit for vacuum is

𝐸𝑠 =𝑚𝑒

2𝑐3

𝑞𝑒ћ≃ 1.3 × 1018 𝑉 𝑚⁄

(5.2)

80

where me is the mass of electron, c is light’s speed in vacuum, qe is the elementary charge, and ћ

is the reduced Planck constant.

When the wire embedded with the nanoheater is exposed to microwaves a voltage builds

up in the wire, which then may cause an electrical breakdown by a corona discharge. Since the

breakdown occurs at the breakdown potential E, the voltage developed between the ends of the

wire V is given by

𝑉 = 𝐸 × 𝑟 (5.3)

where E can be the breakdown potential for air or argon gas due to Corona discharge, or the

Schwinger limit for vacuum, and r is the tip radius of the wire. The value of r depends to the

diameter of wire as well as the cutting method used, and the largest r could be half of the

diameter. So for a wire with diameter of 0.05 mm:

𝑉 = 3 × 106 × 0.025 × 10−3 = 75 V (5.4)

The resistance of a copper wire with 24 mm in length was measured by a voltmeter as 0.5

. Therefore,

𝐽 =𝑉2

𝑅=

752

0.5= 11250 J/s (5.5)

81

Comparing the above calculated Joule energy per unit time with the values of V2/R obtained in

the Joule heating experiments shown in Table 5.4 suggests that the electrical breakdown due to

microwave irradiation induces more than enough powder for the ignition of the nanoheater.

Wire/reactive composites encapsulated in argon/vacuum atmosphere: In order to do experiments

in atmospheres other than air, wires/nanoheaters with embedded copper wires were sealed in a

Pyrex tube with a length of 300 mm, diameter of 6.40 mm and a wall thickness of 1.30 mm.

Argon atmosphere: First, the Pyrex tube was evacuated with a Platinum JB vacuum pump

which was capable of evacuating at rate of 1.5×10-3 m3/s to a minimum pressure of 4×10

-4

mmHg. Then the process was followed by argon backfilling. This process was repeated twice

before closing the tube using a propane torch.

Vacuum atmosphere: The tube was evacuated for 30 minutes with an evacuation rate of

1.5×10-3

m3/s to an estimated pressure of 4×10

-4 mmHg. Then the tube was closed using the

propane torch.

82


Various exposed lengths of wire and diameters were investigated to determine the

window for microwave ignition. The configuration involved having the wire completely straight

while embedding. The nanoheater could be placed in the center of the wire as shown in Figure

5.9, or at any position on the wire.

Figure 5.10 shows the ignition time in air for a 2Al-Fe2O3-3(Al-Ni) with embedded a

0.05 copper wire vs. the wire length. In order to understand the effect of reactive composite on

the ignition time, copper wires with the same diameter and same lengths were subjected to

microwaves to determine the time at which sparking occurred. Figure 5.10 also shows the

sparking time of the free wires. All the data points in Figure 5.10 represent the average values of

four repeated experiments for each condition. It is seen that for the same total length of copper

wire, the ignition time and the sparking time are comparable. Thus, the reactive composite

ignited when the wire sparked. It is also seen that for a total length (l = l1 + l2 + d) less than 24

mm there is no sparking on the copper wire, with or without reactive composite.

Various exposed lengths of aluminum wires with diameter of 0.1 mm were also subjected

to microwave ignition tests in air and the results are presented in Figure 5.11. There was no

sparking for l < 24 mm. Figure 5.11 also shows that by increasing the wire length the ignition

83

time decrease. However, the ignition time increase to 4 s at 60 mm. Further work is needed to

more fully understand the effects of the length of aluminum wire.

Figure 5.10: Ignition time of a 2Al-Fe2O3-3(Al-Ni) composite with a 0.05 mm embedded

copper wire vs. the length of the copper wire. Copper wires with no

composites sparked at times comparable to the ignition times.

20 30 40 50 60 700

2

4

6

8

10

l (mm)

t (s

)

Copper wire embedded in 2Al-Fe2O

3-3(Al-Ni)

Copper wire

84

Figure 5.11: Spark time vs length for aluminum wire in air.

The electrical breakdown field strength (E) is different for different atmospheres. The

microwave experiments were repeated in argon. Figure 5.12 shows the spark time for different

lengths of copper. The results showed that the spark happened in much shorter lengths for an

argon atmosphere than for air. For copper or aluminum wires, no spark was observed in air for

wire lengths less than 24 mm, but for argon the shortest length at which a spark happened is 10

mm.

20 30 40 50 601

2

3

4

5

6

l (mm)

t (s

)

85

Figure 5.12: Spark time vs length for copper wire with diameter of 0.05 mm in argon

atmosphere.

In order to prove that the nanoheaters can be ignited by microwave irradiation in the

absence of a gaseous atmosphere, microwave experiments were repeated in vacuum. As

discussed in Section 5.3.1, the electrical breakdown in vacuum occurs by a different mechanism,

as the electric potential difference induced in the copper wire subjected to microwaves reaches

the Schwinger limit. This work determined that the critical length of a copper wire with a

10 15 20 25 30

5

10

15

20

25

30

35

40

l (mm)

t (s

)

86

diameter of 0.05 mm above which sparking happens was 13 mm, at which a sparking time was

3.90 s. The ignition time was reduced to 2.30 s at a copper wire length of 15 mm.

87

6. Microscale Joining

This chapter explains how the fabricated reactive composites can be used for soldering

copper. A new method for soldering aluminum is also introduced.

6.1 Copper Soldering Using Reactive Composites as Heat

Sources

Fabricated nanoheaters were tested as heat sources for soldering. For this purpose, hybrid

bimetallic nanoheaters were ultrasonically consolidated with two strips of copper which provided

leads for ignition by Joule heating. The consolidation of the nanoheater was done directly on a

sheet of copper, Figure 6.1. The consolidated assembly was then put on a layer of solder, and the

nanoheater was ignited. The heat generated by the nanoheater was conducted through the

thickness of the copper sheet and melted the solder layer, which resulted in good metallutgical

joining of the copper sheet and the solder.

88

Figure 6.1: Consolidating of nanoheater with two copper strips embedded on a copper

sheet.

6.1.1 Preparation of Solder Sheet

Thin sheets (discs) of tin-base lead-free solder (SN100C®) 6 mm in diameter and 0.5 mm

in thickness were prepared by consolidating 50 μm mono-disperse solder balls (Figure 6.2)

produced by the uniform-droplet spray (UDS) process, a capillary jet breakup process that

produces mono-disperse metal balls [94]. Consolidation of the solder balls was done by

ultrasonic powder consolidation (UPC). All solder discs used in this work were consolidated

under the conditions optimized for the solder balls (consolidation temperature: 100 °C,

consolidation time: 3 s, vibration amplitude: 9 μm, uniaxial pressure: 100 MPa). Alternatively,

solder sheets of desired thickness fabricated by any other methods may be used.

89

Figure 6.2: Mono-disperse SN100C® droplets produced by the UDS process.

Figure 6.3 shows a schematic of the joining experiment. First, the consolidated SN®100C

sheet is placed at the bottom. A 1×1×0.05 cm copper sheet is placed on top of the solder sheet.

Then, a 2Al-3CuO-1(Al-Ni) nanoheater consolidated together with two copper strips (leads) is

put on the solder sheet with the copper sheet directly in contact with the solder sheet.

Figure 6.3: Schematic of joining setup of a copper sheet to a solder layer.

90


Figure 6.4 shows the cross section of a 2Al-3CuO-1(Al-Ni) reactive composite

consolidated with two embedded copper strips on top of the copper sheet. The dimensions of the

consolidated reactive composite are in dimensions of 6 ϕ × 0.5 ± 0.1 (mm).

Table 3.1 suggests that the heat generated by this reactive composite per unit volume is

1.8×107 kJ/m

3, so the heat generated by the reactive composite presented in Figure 6.4 is

𝑄 = 1.8 × 107 ×(6 × 10−3)2𝜋

4× 0.5 × 10−3 = 0.25 kJ

(6.1)

Figure 6.4: Cross section of a reactive composite consolidated with two embedded copper

wires on top of a copper sheet.

91

The amount of heat needed to melt a solder layer, Q1, is given by

𝑄1 = 𝑞1 + 𝑞2 (6.2)

where q1 is the heat required to heat the solder to the melting point, which is given by

𝑞1 = 𝑚Sn100C × 𝐶𝑝Sn100C(𝑇𝑚 − 𝑇0) (6.3)

and q2 is the latent heat of fusion of the solder

𝑞2 = 𝑚Sn100C × ℎ𝑓Sn100C (6.4)

where m is the mass of solder layer, 𝐶𝑝SN100C is the heat capacity of SN

®100C (0.22 J /g °C)

[101], ℎ𝑓Sn100C is heat of fusion (60.70 J/ kg) [102] and Tm is the melting point of solder (227 °C)

[103]. The mass of a solder sheet (described in Section 6.1.1: 6 mm in diameter and 0.5 mm in

thickness) is measured as 0.1 g. Thus, the minimum heat for melting an Sn100C solder disc with

the mass of 0.1 g is

𝑄1 = 0.1 × 0.22 × (227 − 21) + 0.1 × 60.70 = 10.60 J (6.5)

To assess the amount of heat required to melt the solder, assume for simplicity that the

copper sheet is also heated uniformly to the same temperature as that of the solder, i.e., 227 °C.

This gives the upper-bound value of the amount of heat required to heat the Cu plate to 227 °C,

Q2, which is given by:

92

𝑄2 = 𝑚𝐶𝑢 × 𝐶𝑝𝐶𝑢(𝑇𝑚

𝑆𝑛100𝐶 − 𝑇0) (6.6)

The mass of the copper sheet (described in Section 6.1.1: 0.1×0.1×0.05 mm) is measured

as 0.4 g. Substituting 𝑚𝐶𝑢 = 0.4 g, 𝐶𝑝𝐶𝑢 = 0.385 (J /g °C) [103] and 𝑇𝑚

𝑆𝑛100𝐶 = 227 °C,

𝑄2 = 31.72 J.

Thus, the minimum heat required to melt (reflow) the solder is

Q1+Q2=10.60+31.72=42.32 J.

In actuality, more heat may be required, as the above calculation ignores the dissipation

due to radiation and conduction. However the use of the 2Al-3CuO-1(Al-Ni) nanoheater with an

energy output of 250 J provided more than enough heat to achieved good reflow while keeping

the copper sheet without melting the copper sheet at the interface with the nanoheater. Figure 6.5

shows the cross section of the reflowed solder layer on the copper sheet. Full melting and

resolidification of the solder layer with good bonding with the copper sheet is apparent.

93

Figure 6.5: Cross section of soldered copper using 2Al-3CuO-1(Al-Ni) reactive composite.

6.2 Flux-Less Soldering of Aluminum

This section presents an economical and dependable method for flux-less direct soldering

of aluminum with tin-base solders in which use of flux is replaced with in-process ultrasonic

abrasion to activate the aluminum surface for fresh metal-to-metal contact with solder. This

invention is presented in an international patent application [105] and in a paper by the author: S.

Gheybi-Hashmebad, Z. Gu and T. Ando, “Flux-less direct soldering of aluminum by ultrasonic

surface activation,” Journal of Materials Processing Technology, (under review as of December

2015).

94


Process concept: The flux-less soldering process investigated in this section employs fine

abrasive corundum alumina particles to activate aluminum surface. A stack of aluminum and tin-

base solder sheets sandwiching a very thin layer of fine alumina particles, held under moderate

clamping pressure (~20 MPa), Figure 6.6(a), is first subjected to in-plane ultrasonic vibration at

room temperature for a brief duration (~1 s) to disrupt the oxide on the aluminum surface and

simultaneously allow for direct aluminum-solder contact, Figure 6.6(b). This process also purges

residual oxygen from the solder-aluminum interface. The stack is immediately heated for reflow

while keeping the stack under the clamping pressure, Figure 6.6(c). Upon solidifying the molten

solder, a well-bonded solder-aluminum interface is obtained. This process requires no fluxing

and is done in open air. The procedures of the experiment performed are elaborated below.

Figure 6.6: Ultrasonic abrasive activation of aluminum surface.

95

Sprinkling solder sheet with abrasive powder: The preparation of solder discs are explained in

6.1.1. In this section the method for coating solder sheet with abrasive powder is explained. A

suspension-sedimentation method was used to sprinkle fine alumina powder over the solder

discs, Figure 6.7. The solder discs were first placed in beakers 50 mm in diameter. Suspensions

of fine alumina powder in ethanol were prepared in different alumina densities by rotary mixing

in a 40 mm-diameter cylindrical container rotated horizontally at 750 rpm for 5 minutes,

followed by ultrasonic sonication for 10 minutes and additional rotary mixing for 20 minutes.

Alumina powders of two different nominal particle sizes, 0.3 μm and 1.0 μm, purchased from

Mark V Laboratory, were used. The alumina-ethanol suspensions were slowly poured in separate

beakers in which the solder discs were laid. The suspensions in the beakers were kept still to

allow the alumina particles to sediment while ethanol evaporated. After complete evaporation of

ethanol, the solder discs were left on the bottom of the beakers sprinkled with a desired amount

of alumina powder. The concentration of alumina sprinkled on the solder disc, x, defined as the

mass per unit area, was calculated from x = M / πr2 where M is the total mass of alumina powder

in the suspension and r is the radius of the beaker. In the present work, M was adjusted to

prepare solder discs sprinkled with 0.3 μm alumina at x = 0, 1, 2, 3 and 4 g/m2. To provide

comparison, solder discs sprinkled with 1.0 μm alumina were prepared at x = 1 and 2 g/m2.

96

Figure 6.7: Sedimentation sprinkling of corundum particles on solder discs.

Ultrasonic abrasive surface activation and reflow: Two types of specimens were fabricated, one

for microscopic characterization of the solder-aluminum interface and the other for evaluating

the strength of solder-joined aluminum sheets by tensile shear testing. To fabricate a specimen

for microscopic characterization, an aluminum sheet (1 mm thick 1100 Al sheet from McMaster-

Carr) was placed on the heater plate of the ultrasonic welder, and an alumina-sprinkled solder

disc was put on the aluminum sheet with the alumina side of the solder disc facing the aluminum

sheet as illustrated in Figure 6.6(a). Then a punch, 5 mm thick and 6 mm in diameter, made of

stainless steel, was placed on top of the stack of solder disc and aluminum sheet. The punch was

needed to avoid direct contact of the sonotrode tip to the soft solder disc. No abrasive powder

was applied between the stainless steel punch and the solder disc. To fabricate tensile specimens,

two alumina-sprinkled solder discs stacked with their alumina-free faces mating were

sandwiched with two rectangular aluminum sheets and placed on the base plate of the ultrasonic

welder as illustrated in Figure 6.8(a). The two aluminum sheets, prepared in the dimensions 10 ×

97

15 × 1 mm, had a 3 mm through pin hole 5 mm from one of the ends to provide a grip on a

soldered specimen for tensile loading. Then, for either type of specimen, the sonotrode was

brought down to clamp the stack under a normal pressure of 20 MPa and subject the stack to 20

kHz in-plane ultrasonic vibration for 1 s at room temperature at an amplitude of 9 μm.

Figure 6.8: Schematic of fabrication of an aluminum-solder-aluminum tensile specimen (a)

ultrasonic surface activation, (b) solder reflow.

98

Immediately after the application of ultrasonic vibration, the heater plate was turned on to

bring the temperature of the specimen, still under 20 MPa, above the liquidus of SN100C® (240

°C [94]) for solder reflow, Figure 6.8(b). Figure 6.9 shows the profile of heater temperature over

the periods of heating, reflowing and cooling of an aluminum-solder specimen. Similar schedules

are normally used for SN100C® [95]. The specimen reached the liquidus temperature of the

solder in 90 s, after which reflow occurred over a controlled period of 45 s. The same reflow

schedule was used for all other specimens.

Figure 6.9: Solder reflow thermal profile.

Microscopic evaluation and tensile testing: The UPC-consolidated solder discs and the reflowed

specimens were examined under an Olympus VANOX-T optical microscope using standard

99

metallographic specimen preparation procedure. All specimens were etched with 2% nital. The

bond strength of the solder-joined aluminum specimens was tested on a 2710 -103 Instron testing

machine with a maximum load of 1 kN at a constant crosshead speed of 0.5 mm/min in the

manner schematically illustrated in Figure 6.10. The tensile specimens were activated with 0.3

μm alumina particles at x = 1, 2, 3 and 4 g/m2 for 1 s at room temperature and reflowed in the

manner shown in Figure 6.9. The tensile force, F, vs. the imposed displacement, d, was recorded

using a load cell that had a resolution of ±2.5mN. The stress on the specimen was calculated as

F / πr2

where r is the radius of the solder disc. The maximum value of F / πr2

on the stress-

displacement curve was regarded as the joint strength of the specimen.

Figure 6.10: Schematic of tensile shear test.

100


Reflowed specimens: Figure 6.11(a) shows an optical micrograph of a cross section of an as-

consolidated solder (SN100C®) disc. The original mono-sized solder balls are seen to have been

deformed and consolidated into a dense material. The prior solder particles, delineated with

oxide films, exhibit a fine cellular/dendritic microstructure from the rapid solidification they

underwent. Subjecting the alumina-activated specimens to reflow under the schedule in Figure

6.11(b) caused melting and re-solidification of the solder. The reflowed solder layer exhibits a

cellular/dendritic microstructure, Figure 6.11(b), which represents the solder microstructure in all

reflowed specimens. The cells/dendrites are seen to have grown through the prior particle

boundaries of the consolidated solder disc, which verifies the melting of the solder during the

reflow under the applied schedule.

101

Figure 6.11: Cross sections of etched solder layer for (a) as consolidated, (b) after reflow.

102

Figure 6.12 shows as-polished cross sections of reflowed solder-aluminum specimens

that were activated with 0.3 μm alumina at x = 1 - 4 g/m2. The specimen prepared at x = 1 g/m

2,

Figure 6.12(a), exhibits gaps at the interface where bonding was not achieved. SEM confirmed

the presence of gaps, Figure 6.13(a). Increasing alumina concentration to 2 g/m2

drastically

improved the bonding as seen in Figure 6.12(b) where the interface is largely clean and free of

voids or gaps, except for very small amounts of alumina particles remaining at the interface.

SEM confirmed good bonding at the interface, Figure 6.13(b). Good interfacial bonding was

maintained when alumina concentration was further increased to 3 g/m2, Figure 9(c), although

somewhat increased amounts of alumina particles remained at the interface, forming small,

occasional colonies at the interface. More alumina colonies were left at the interface in the

specimen prepared at 4 g/m2, Figure 6.12(d), leaving less interfacial area available for bonding.

High-resolution SEM reveals the alumina particle colonies more clearly, Figure 6.13(c). Thus,

with 0.3 μm alumina, an optimum range of x is found to be 2 - 3 g/m2. Above the optimum

alumina concentration range, alumina colonies remaining at the interface restricted the wetting of

activated aluminum surface with the molten solder during the reflow, whereas below the

optimum range surface activation was insufficient. In fact, the interface of a control specimen,

prepared with no alumina but under otherwise identical conditions, showed no signs of bonding,

Figure 6.14.

103

Figure 6.12: Cross section of soldered aluminum activated with 0.3 μm alumina at (a) 1

g/m2, (b) 2 g/m

2, (c) 3 g/m

2, (d) 4 g/m

2,

104

Figure 6.13: SEM micrographs of the cross section of aluminum-solder specimens

activated with 0.3 μm alumina: (a) gaps at interface (x = 1 g/m2), (b) well

bonded interface (x = 2 g/m2), (c) alumina colonies remaining at interface (x

= 4 g/m2).

105

Figure 6.14: Cross section of a control aluminum-solder specimen reflowed without

surface activation.

Effective surface activation was obtained in specimens prepared with 1 μm alumina as

well, but at a lower alumina concentration of 1 g/m2, Figure 6.15(a). At x = 2 g/m

2, nearly half of

the interface was associated with elongated colonies of alumina particles, Figure 6.15(b). Thus,

the use of 1 μm alumina shifted the optimum alumina concentration to lower values and in a

narrower range. This suggests that coarser alumina particles produce stronger abrasion effects

and hence good bonding at lower alumina concentration, but their colonies create a stronger

barrier to the wetting between aluminum and molten solder, hence the lower and restricted

optimum alumina concentrations.

106

Figure 6.15: Cross sections of soldered aluminum activated with 1 μm alumina at (a) 1

g/m2 and (b) 2 g/m

2.

To quantitatively address the extent of metallurgical bonding achieved in the solder-

aluminum specimens, the fraction of the interface joined, f j , was calculated with f j = l j / lo

where l j is the length of the part of interface in a cross section that is judged joined and lo is the

length of interface. Figure 6.16 plots f j against the alumina concentration x for the specimens

activated with 0.3 μm and 1 μm alumina. It is clear that the effect of abrasive surface activation

with alumina powder depends on both the size and concentration of alumina particles used, and

that the optimum concentration range is wider and at higher concentrations when activation is

done with finer alumina particles.

107

Figure 6.16: Fraction of the interface joined of aluminum-solder-aluminum specimens vs.

concentration of alumina used for surface activation.

Joint strength: The joint strength of aluminum sheets solder-joined with abrasive surface

activation was examined by the tensile shear test shown in Figure 6.10 using tensile specimens

activated with 0.3 μm alumina at concentrations 1, 2, 3 and 4 g/m2

joined in the manner

illustrated in Figure 6.8. Figure 6.17 shows a stress - displacement curve obtained with a

specimen activated at an alumina concentration of 2 g/m2. The joint strength, defined as the

stress at fracture, is 48 MPa for this specimen which is indicative of satisfactory bonding as it

compares well with typical values of Sn100C and other Sn-base solder joints [96- 98]. A similar

joint strength of 45 MPa was obtained with the specimen prepared at 3 g/m2. However, much

lower strengths were determined at 1 and 4 g/m2. Figure 6.18 plots the joint strengths of the four

0 1 2 3 4 50

0.5

1

Alumina concentration

Fra

cti

on

of

inte

rfa

ce

jo

ine

d

0.3 m

1 m

108

tensile specimens normalized by the highest value obtained (48 MPa obtained at 2 g/m2) against

x, which is virtually identical to the plot of the fraction of interface joined in Figure 6.16. Thus,

the optimum alumina concentrations determined by the tensile tests of the aluminum-solder-

aluminum specimens coincide with those that were determined microscopically for the solder-

aluminum specimens.

Figure 6.17: Stress - displacement curve of an aluminum-solder-aluminum specimen

activated with 0.3 μm alumina at 2 g/m2.

0 0.2 0.4 0.6 0.8 10

10

20

30

40

50

Distance (mm)

Str

ess (

MP

a)

109

Figure 6.18: Normalized Joint strength of aluminum-solder-aluminum specimens vs.

concentration of 0.3 μm alumina used for surface activation.

The good metallurgical bonding in an optimally joined aluminum-solder-aluminum

specimen is further confirmed in Figure 6.19 which shows a cross section of a specimen

activated with 0.3 μm alumina at an optimum alumina concentration of 2 g/m2. Both the top and

bottom interfaces in the specimen are clean and exhibit no voids or gaps, which attests to the

good metallurgical bonding of the solder layer to the aluminum sheets. The aluminum-solder-

aluminum tensile specimens joined under the optimum conditions all exhibited a fracture surface

indicative of fracture propagation through the solder layer and not along the aluminum-solder

interface. Figure 6.20 shows such a fracture surface of a specimen activated 0.3 μm alumina at x

= 2 g/m2 where interdendritic fracture through the solder layer is apparent. No evidence was

0 1 2 3 4 50

0.5

1

Alumina concentration (g/m2

)

No

rma

lize

d j

oin

t str

en

gth

Maximum joint strength= 48 MPa

110

found for crack initiation or propagation caused by alumina particles at the interface, indicating

that the alumina particles left in small amounts at the interface were benign.

Figure 6.19: Cross section of aluminum sheets solder-joined with ultrasonic abrasive

surface activation using 0.3 μm alumina at x = 1 g/m2.

111

Figure 6.20: A specimen activated with 0.3 μm alumina at x = 2 g/m2 shows fracture path

through the solder layer.

6.2.3 Aluminum Flux-Less Soldering Using Reactive Composites

Instead of using a continuous reflow process as in Figure 6.9, a reactive composite was

tested as a heat source for reflow in the flux-less soldering of aluminum developed in the present

study. Combining local heating with such a heat source with the flux-less soldering process

potentially removes the need for oven heating for reflow, which would lead to increase use of

soldering at reduced cost in manufacturing.

112

The reactive composite tested for this purpose was a 2Al-3CuO-2(Al-Ni) nanoheater. It

was consolidated in the same manner as illustrated in Figure 6.1. A disc of solder (Sn100C®),

consolidated from mono-sized powder by UPC in the same manner described in section 6.1, was

used. Figure 6.21 schematically illustrates the application of nanoheater in flux-less soldering of

aluminum tested in the present work while different configurations of nanoheater-aluminum-

solder stacking are possible. In this study, the solder layer was placed at the bottom on the heater

plate, with its side sprinkled with alumina particles facing up, 0.3 μm alumina particles was

applied at an optimum condition of 2 g/m2 as illustrated in Figure 6.12. On top of the aluminum-

sprinkled solder layer, a sheet of aluminum, 1 mm in thickness, was placed. On top of the

aluminum sheet, a 2Al-3CuO-2(Al-Ni) nanoheater consolidated with a pair of thin (0.1 mm)

copper wires was placed. Then, the whole stack was clamped under a uniaxial process of 10 MPa

and subjected to in-plane ultrasonic vibration (20 kHz, 9 μm amplitude) for 1 s, to activate the

surface of the aluminum sheet facing the solder layer and purge oxidation from the aluminum-

solder interface.

Figure 6.21: Schematic of joining setup of a copper sheet to a solder layer.

113

The nanoheater was then ignited by applying a voltage between the two copper wires

which caused joule heating in the nanoheater. The amount of heat oupput of the nanoheater was

calculated from the heater mass to be 0.15 kJ, which was sufficient to conduct enough heat

through the aluminum sheet to melt the solder while keeping the aluminum sheet largely

unaffected by the heat. In commercial implementations an intermediate conduction layer may be

used between the nanoheater and the aluminum part. The mass, and hence the heat output of the

nanoheater, should also be optimized for the thickness of aluminum parts to be soldered as well.

Figure 6.22 shows the cross section of a specimen soldered in the manner described above.

Despite the very short duration of the heating with the nanoheater (which is estimate to be 1~2

s), the interface is free of voids or gaps, exhibiting good metallurgical joining.

Figure 6.22: Cross section of soldered aluminum activated with 0.3 μm alumina at 2 g/m2.

114

7. Conclusions

1. Ultrasonic powder consolidation (UPC) was used to fabricate new types of hybrid reactive

composites, 2Al-Fe2O3-x(Al-Ni) and 2Al-3CuO-x(Al-Ni), which combine the low-

temperature ignitability of Al-Ni composites and the high heat output of thermites in single

composites. The major findings use described in the bullets below:

Reactive composites 2Al-Fe2O3-x(Al-Ni), 2Al-3CuO-x(Al-Ni) and 2Al-Al2O3-x(Al-Ni), (x

= 1 - 4), ultrasonically consolidated from nano-thick flakes of Al and Ni mixed with Fe2O3,

CuO or Al2O3 nanoparticles at 573 K under 100 MPa pressure, exhibited a composite

structure in which Ni flakes and oxide particles were distributed evenly in a fully densified,

metallurgically bonded Al matrix.

During continuous heating at 125 K/min (2.08 K/s), the composites began self-heating

before they ignited. The self-heating was caused by solid-state exothermic reactions

between Al and Ni which accelerated just before the ignition. The acceleration stage

occurred at higher temperatures in composites with lower x since they had more oxides that

retarded the Al-Ni reactions at the An-Ni interface. The Fe2O3 and CuO nanoparticles,

being clustered in the composites, had stronger retarding effects than the Al2O3 particles. In

115

all composites, ignition was triggered by liquid formation due to eutectic melting and/or

melting of aluminum which was noticed by the occurrence of endothermic shifts on

thermographs. The 2Al-Fe2O3-1(Al-Ni) and 2Al-3CuO-1(Al-Ni) composites ignited when

Tsp reached the melting point of aluminum, but all other composites ignited below the

melting point of aluminum, indicating that the Al-Ni reaction fronts in the composites had

reached the eutectic temperature or the melting point of aluminum.

The ignition of the 2Al-Fe2O3-x(Al-Ni) and 2Al-3CuO-x(Al-Ni) composites by the thermite

reaction occurred in two stages: an initial localized stage where the liquid produced by the

Al-Ni reactions reacted with the oxide, and a global stage where the molten aluminum

produced in the initial stage reacted extensively with the oxide. The ignition of the 2Al-

Al2O3-x(Al-Ni) composites occurred more gradually as the Al-Ni reactions produced liquid

over time. The heating rates just after ignition indicated that the thermite reactions were

about two orders of magnitude faster than the Al-Ni reactions.

The hybrid bimetallic-thermite compositions, 2Al-Fe2O3-x(Al-Ni) and 2Al-3CuO-x(Al-Ni),

combine the low-temperature ignitability of Al-Ni composites and the high heat output of

thermites in single composites.

The temperature increases due to solid-state exothermic reaction between Al and Ni in the

hybrid nanoheaters were simulated with a model based on a non-isothermal Avrami

equation.

116

2. Different methods for igniting reactive composites are introduced in this study including:

spark ignition, continuous-heating ignition, Joule heating ignition, and microwave ignition.

The latter two techniques established in this work, provide easy and economical techniques

for igniting nanoheaters that are adaptable to industrial microheating processes in any

environment.

3. Nanoheaters developed in this work find applications in micro-volume heating, microjoining

in particular, where effective means for local heating are required. The effectiveness of the

nanoheaters was confirmed in the soldering of aluminum and copper. The application of

nanoheater to the flux-less soldering of aluminum was enabled also by implementing an

ultrasonic surface activation technique developed in this study. In the ultrasonic surface

activation process, alumina particles are used for removing the oxide layer on the surface of

aluminum. Through the systematic experiments alumina concentration and size were

optimized for joint strength comparable to typical values of tin-base solder joints were

achieved.

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8. Recommended Future Work

Additional work is recommended in two areas.

First, a comprehensive thermal analysis is needed in order to take full advantage of the

micro-volume heating with reactive composites (nanoheaters). This is because the amount of

heat generated by the nanoheater may not entirely be exploited to affect the intended local

heating due to the possible dissipation by radiation, evaporation and mass expulsion which takes

away part of the reaction enthalpy, reducing the net amount of heat available for local heating.

The heat dissipation due to these processes generally increases with increasing reaction rate and

adiabatic temperature of the nanoheater, and as such is a concern with the hybrid nanoheaters

developed. In fact, it was noticed in the present work that hybrid nanoheates with small amounts

of bimetallic additions, e.g., x ≤ 1, reacted explosively causing expulsion of the hot reaction

products, whereas the reacted hybrid nanoheaters with higher bimetallic additions stayed intact.

To determine the extent of heat dissipation, and hence the amount of net enthalpy for

local heating, it is recommended to (1) verify the theoretical reaction enthalpy given in Table 3.1

and (2) determine the amount of heat that effects the local heating intended. An effective method

to achieve (1) is differential scanning calorimetry (DSC). (2) may be achieved by an experiment

where a nanoheater placed on a metal substrate into which the reaction enthalpy is conducted is

118

ignited while monitoring the temperature distributions in the substrate over time. The metal

substrate may be strip (1 D), sheet (2D) or plate (3D). Then, the heat absorbed by the metallic

substrate may be determined using CFD methods. This gives the upper bound value of the ‘net’

heat available for the micro-volume heating of concern.

While knowing how much of the reaction enthalpy contributes to the intended local

heating is essential, avoiding an expulsive reaction would promote effective use of reaction

enthalpy. This can be achieved by (1) optimizing the nanoheater compositions and structure and

(2) devising ways to incorporate nanoheaters within the material to be heated so that there is no

path for the dissipation of unused heat. To this end, simultaneous ultrasonic consolidation of

nanoheaters in the material to be local-heated should be investigated.

Another area where extensive future work is needed is the industrial implementation of

local heating with nanoheaters. While additional tuning and optimization of the nenoheater

composition and structure are still required, the most critical issue in commercialization is the

development of economical means for the high-rate production of nano-structured reactive

composites. The results of this work strongly indicates that UPC is more advantageous than other

methods, such as powder metallurgy, PVD and CVD processes, and would be viable for

commercialization if it is made adaptable to high-rate production. The latter is considered

feasible via R & D based on roll-type ultrasonic sheet welders that are already available

commercially. The key issues in the required R & D are the development of an ability to

119

consolidate sheet of powder compact at elevated-temperature while applying ultrasonic vibration

through the roller.

120

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133

Appendix: MATLAB Codes

I. Al-Ni interface length calculation:

[m n]=size(Ni_boundary);

A=zeros(m,n);

for ii=1:m

for jj=1:n

if Al_boundary (ii,jj)==1 && Ni_boundary (ii,jj)==1

A(ii,jj)=1;

end

end

end

imshow(A)

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II. Interfacial area for unit volume of a reactive composite:

[m n]=size(A);

m=0:10:m;

c=[];

d=[];

for mm=0:10:m

clear c

c=[];

for ii=m+1:m+10

for jj=1:n

if A(ii,jj)==1

c=[c,1];

end

end

end

[l n]=size(c);

d=[d,n];

end

S_layer=mean(d)

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III. Kinetic model:

clc, clear %inputs T_0 = 500; %initial temperature (K) S_i = 3e6; %0.49*2e7; %initial interfacial area (m^2/m^3) % W = 2.5*10^-7; %compound thickness (m) H_C = 150624; %enthalpy of formation at 298 K (J/mol) K = 10000; %lateral growth factor (m/s) k = 8.617343*10^-5; %Boltzmann constant (eV/K) Q = 1.44; %lateral growth activation energy (eV) N = 10^16; %number of nuclei per unit area (1/m^2) V_C = 2.4*10^-5; %molar volume of the compound phase (m^3/mol) D_0_Ni_Al=1.77e-7; %pre exponential factor for diffusivity Ni to Al D_0_Al_Ni=8.19e-8; %pre exponential factor for diffusivity Al to Ni R=8.3144621; %Gas constant E_A_Ni_Al=43.65e3; % activation energy Ni to Al E_A_Al_Ni=34.7e3; % activation energy Ni to Al dt = 0.00024; %time step (s) steps=10000; %number of iterations g = 9.81; L_c = 10^-5; A_s = 10^-4;

for ii = 1:steps if ii==1 T = T_0; else T = T1(ii-1); end G(ii) = K*exp(-Q/(k*T)); %lateral growth rate G1(ii) = sum(G); D_Ni_Al(ii)=1.77e-7*exp(-43.65e3/(R*T)); D_Al_Ni(ii)=8.19e-8*exp(-34.7e3/(R*T)); W(ii)=D_Ni_Al(ii)/G1(ii)+D_Al_Ni(ii)/G1(ii); x_a(ii) = 1-exp(-N*pi*(dt^2)*(G1(ii)^2)); x_v(ii) = S_i * W(ii) * x_a(ii); %volume fraction if ii==1 dx_v(ii)=x_v(ii); else dx_v(ii)=x_v(ii)-x_v(ii-1); end %heat loss

136

T_f(ii) = (T+T_0)/2; beta(ii)=1/T_f(ii); Ra(ii) = (g*beta(ii)*(T-T_0)*(L_c^3)*Pr(T))/(Nu(T)^2); Nus(ii) = 0.54 * (Ra(ii)^0.25); h(ii) = k_air(T) * Nus(ii) / L_c; delH_L(ii) = h(ii) * A_s * dt * (T-T_0); %thermal balance h1(ii)=quad(@Cp_Al_mol,298,T); h2(ii)=quad(@Cp_Ni_mol,298,T); delH_1(ii)=3*h1(ii) + h2(ii); delH_2=H_C; delH_3(ii)=quad(@Cp_Al3Ni_mol,298,T); delH_4(ii)=quad(@Cp_Fe2O3_mol,298,T); delH(ii)=-delH_1(ii)-delH_4(ii)+delH_2+delH_3(ii); delH_V(ii)=delH(ii)/V_C; x_A(ii)=0.5-(0.86*x_v(ii)); x_B(ii)=0.2-(0.19*x_v(ii)); x_C=0.3; B(ii) =0.8* [(delH_V(ii) * dx_v(ii)*A_s*L_c)-delH_L(ii)]; sm = 0; dT = 0.0001; if B(ii)<0 while sm < abs(B(ii))

Cp=0.2*((x_A(ii)*Cp_Al(T)*A_s*L_c)+(x_B(ii)*Cp_Ni(T)*A_s*L_c)+(x_C*Cp_F2O3(T)*A_s*L_c)+(x_v(ii)*C

p_Al3Ni(T)*A_s*L_c)); sm = sm + Cp*dT; T=T-dT; end end if B(ii)>0 while sm < B(ii)

Cp=0.2*((x_A(ii)*Cp_Al(T)*A_s*L_c)+(x_B(ii)*Cp_Ni(T)*A_s*L_c)+(x_C*Cp_F2O3(T)*A_s*L_c)+(x_v(ii)*C

p_Al3Ni(T)*A_s*L_c)); sm = sm + Cp*dT; T=T+dT; end end if B(ii)==0 T=T; end

t(ii) = ii*dt;

137

T1(ii) = T; if x_a(ii) > 1 break end if x_v(ii) > 1 break end if T1(ii) > 913 break end end plot(t,T1,'LineWidth', 2) ylabel('Temperature (K)', 'FontSize', 12) xlabel('Time (s)', 'FontSize', 12)

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