Characterization of Low Melting Temperature, Low … · Characterization of Low Melting...

Characterization of Low Melting Temperature, Low-Ag,

Bi-Containing, Pb-Free Solder Alloys

by

Eva Kosiba

A thesis submitted in conformity with the requirements for the degree of Masters of Applied Science

Materials Science and Engineering University of Toronto

© Copyright by Eva Kosiba 2016

ii

Characterization of Low Melting Temperature, Low-Ag, Bi-

Containing, Pb-Free Solder Alloys

Eva Kosiba

Masters of Applied Science

Materials Science and Engineering

University of Toronto

2016

Abstract

Restrictions of lead in solder lead to adoption of SAC305 in consumer products. While

high reliability applications use SnPb, supply constraints are driving the adoption of a

replacement. SAC305 has reliability concerns related to elevated process temperatures

and the formation of Ag3Sn.

Reliability performance of three low-Ag, Bi-containing, low melting temperature alloys

were compared to SAC305. All three alloys under test performed as well or better for

consumer applications. Drop testing and accelerated thermal cycling revealed no

differences that would preclude use of these alloys in production. They allow for the use

of lower Tg printed wire boards materials, which have been shown reliable.

These alloys show promise for high reliability applications. In accelerated thermal

cycling, all alloys outperformed the circuit boards. Bi precipitation resulted in less

degradation to the bulk microstructure. Bi did not impact the IMC formation or growth, a

small amount of Ag mitigated growth of Cu3Sn.

iii

Acknowledgments

I would like to express my gratitude to Professor Doug Perovic, Dr Polina Snugovsky

and John McMahon for their guidance, generosity and support. Their knowledge of

various aspect of this research was an inspiration.

I would also like to thank Lorna Devereux and Marianne Romansky for reading through

the many revisions and offering clarity to my often confusing prose. The following

people provided much laboratory assistance and support for which I am grateful: Russell

Brush, Subramaniam Suthakaran, Michael Thomson, Hissan Syed, André Delhaise and

Salvatore Boccia. I would also like to thank Professor Erb and Professor Wang for

agreeing to participate on the review committee on short notice.

Finally I would like to acknowledge the support of the Quality and Reliability Laboratory

at Celestica and the Department of Materials Science and Engineering at the University

of Toronto.

iv

Table of Contents

Abstract… ................................................................................................................... ii Acknowledgments ..................................................................................................... iii List of Tables ............................................................................................................ vii

List of Figures ............................................................................................................ ix List of Acronyms ..................................................................................................... xiii

Chapter 1 Background ................................................................................................ 1 1 Introduction ................................................................................................. 1

2 Need for New Low Melt Solder.................................................................. 2

2.1 Legislative Changes .................................................................................... 2

2.2 SAC305 and Current Issues with Existing Alloys ...................................... 3 2.3 Search for Replacement Solder ................................................................... 6 2.3.1 Binary Alloys .............................................................................................. 6 2.3.2 Ternary and higher order alloy options ....................................................... 9

2.4 Celestica’s Low Melt Solder Research Program ...................................... 12 2.4.1 Phase 1: Alloy Selection ........................................................................... 13

2.4.1.1 Requirement Formulation ......................................................................... 14 2.4.1.2 Literature Search and Phase Diagram Analysis

6 ...................................... 15

2.4.1.3 Preliminary Alloy Selection6 .................................................................... 20

2.4.1.4 Metallurgical Analysis: DSC and Microstructural Evaluation6 ................ 21

2.4.2 Phase 2: Manufacturing Feasibility Study ................................................ 25

2.4.3 Phase 3: Screening Experiments ............................................................... 25

2.4.3.1 Aerospace and Defense, ............................................................................ 26

2.4.3.1.1 Screening Experiment Structure ............................................................... 27 2.4.3.1.2 Microstructural Assessment ...................................................................... 28

2.4.3.1.3 Accelerated Thermal Cycling (ATC)........................................................ 29 2.4.3.1.4 Vibration ................................................................................................... 33 2.4.3.1.5 Summary of Findings ................................................................................ 34

3 Objective of Thesis: Low Melt Solders for Consumer Sector .................. 34 4 References ................................................................................................. 37

Chapter 2 Solder Joints after Reflow (As Manufactured) ........................................ 41

1 Introduction ............................................................................................... 41 2 Bulk Solder Microstructure – Solidification Process................................ 41

2.1 Cu Dissolution in Molten Solder .............................................................. 41 2.2 Bulk Solder Solidification......................................................................... 43 2.3 Undercooling During Solidification ......................................................... 45 2.4 Formation of Facetted IMCs in Bulk Solder............................................. 47 2.5 Bi in Solution and as a Precipitate ............................................................ 48

3 Sn-Cu Reaction and Formation of Interfacial IMC .................................. 49 4 Cu-Ni-Sn Interface .................................................................................... 53

v

5 Experimental Set Up ................................................................................. 54

5.1 Test Vehicle Assembly ............................................................................. 54 5.2 Test Matrix ................................................................................................ 55 5.3 Test Method .............................................................................................. 56

6 Microstructural Evaluation ....................................................................... 56 6.1 Comparison of Bulk Microstructure ......................................................... 56 6.1.1 Bulk Microstructure of QFP Solder Joints................................................ 57 6.1.2 Bulk Microstructure of BGAs ................................................................... 61 6.2 Comparison of Interfacial IMC Layers ..................................................... 67

7 Summary of Findings ................................................................................ 73 8 References ................................................................................................. 74

Chapter 3 Accelerated Thermal Cycling................................................................... 77

1 Accelerated Testing for Reliability Analysis ............................................ 77 2 Microstructural Evolution ......................................................................... 82

2.1 Changes to Bulk Solder ............................................................................ 82 2.2 Changes to Interfacial IMC during Accelerated Thermal Cycling ........... 85

2.3 Effects of Bi .............................................................................................. 86 3 Experimental Setup ................................................................................... 87 3.1 Materials ................................................................................................... 88

3.2 Test Vehicle .............................................................................................. 89 3.3 Test Strategy ............................................................................................. 90

3.3.1 0-100C Thermal Cycling ......................................................................... 91

3.3.2 Harsh Environment (-55C to 125C) Thermal Cycling .......................... 92

3.4 Post ATC Evaluation and Failure Analysis .............................................. 93 4 Results ....................................................................................................... 94

4.1 Reliability and Failure Analysis Results ................................................... 94

4.1.1 0C to 100C Accelerated Thermal Cycling ............................................. 94

4.1.2 -55C to 125C Accelerated Thermal Cycling ....................................... 101

4.2 Microstructure Evaluation ...................................................................... 109 4.2.1 Bulk Microstructure ................................................................................ 109

4.2.2 IMC Growth during Thermal Cycling .................................................... 111 5 Summary of Findings and Conclusions .................................................. 122 5.1 Findings Based on Reliability Data ........................................................ 122 5.2 Findings Based on Microstructural Observations ................................... 122 6 References ............................................................................................... 124

Chapter 4 Tin Whisker Testing ............................................................................... 126

1 Introduction ............................................................................................. 126 1.1 Whisker Growth Kinetics ....................................................................... 128 1.2 Sources of Compressive Stress ............................................................... 130 1.3 Morphology of Sn Whiskers ................................................................... 132 1.4 The Effects of Bi in Solder on Whisker Formation ................................ 133

2 Experimental Set Up ............................................................................... 135 2.1 Materials ................................................................................................. 136

vi

2.2 High Temperature High Humidity .......................................................... 136

2.3 Thermal Shock ........................................................................................ 137 2.4 Post Exposure Evaluation ....................................................................... 138 3 Results ..................................................................................................... 139

3.1 High Temperature High Humidity Results ............................................. 139 3.2 Thermal Shock Results ........................................................................... 139 4 Summary of Findings and Conclusions .................................................. 152 5 References ............................................................................................... 154

Chapter 5 Mechanical (Drop) Shock Testing ......................................................... 156 1 Introduction ............................................................................................. 156 2 Experimental Set Up ............................................................................... 157 2.1 Materials ................................................................................................. 158

2.2 Test Vehicle ............................................................................................ 158 2.3 Assembly................................................................................................. 160

2.4 Test Strategy ........................................................................................... 160 3 Reliability Results ................................................................................... 164

4 Failure Analysis and Microstructural Evaluation ................................... 168 4.1 Dye and Pry Procedure ........................................................................... 169 4.2 Failure Isolation Procedure ..................................................................... 169

4.3 Evaluation of High Tg Board after Drop Testing .................................... 169 4.4 Evaluation of Normal Tg Board after Drop Testing ............................... 173

5 Summary of Findings and Conclusion .................................................... 175 6 References ............................................................................................... 177

Chapter 6 Summary of Findings and Conclusions ................................................. 178

Chapter 7 Future Work ........................................................................................... 182 1 Celestica/Indium Sponsored Whisker Resistant Solder Paste ................ 182

2 ReMap M1: Lower Temperature Soldering Alloys with Improved

Mechanical and Thermal Fatigue Reliability .......................................................... 184

3 ReMap M2: Alloys, Board and Component Surface Finish Interactions

with Reduced Propensity for Whisker Growth ....................................................... 184

4 ReMap M3: Aging Effect of New Lead-Free Materials on Reliability .. 186 5 References ............................................................................................... 190

vii

List of Tables

Table 1: Sn-based eutectic alloys ........................................................................................ 8 Table 2: Preliminary alloy selection

10,6............................................................................. 21

Table 3: DSC analysis of proposed alloys with 75% SAC305 ......................................... 23 Table 4: Alloys for screening experiments ....................................................................... 26 Table 5: Latin square of paste and finish, and board materials ........................................ 27 Table 6: Build matrix showing number of assemblies...................................................... 28 Table 7: Resultant BGA interconnect composition .......................................................... 29

Table 8: ATC failures on 170C Tg board material .......................................................... 30

Table 9: Reliability results on 170C Tg boards after ATC .............................................. 31

Table 10: ATC failures on 150C Tg board material ........................................................ 31

Table 11: Reliability results on 150C Tg boards after ATC ............................................ 32 Table 12: Vibration failures after 2G testing .................................................................... 33

Table 13: Vibration failures after 5G testing .................................................................... 34 Table 14: Consumer alloys under test ............................................................................... 36

Table 15: Build matrix for as assembled analysis ............................................................ 55 Table 16: Table of components evaluated by cross sectioning......................................... 55 Table 17: Composition of QFPs (wt%) ............................................................................ 57

Table 18: Experimental composition of BGA solder joint ............................................... 62 Table 19: Theoretical composition of BGA solder joint .................................................. 62

Table 20: Results of ANOVA test for equal variance and compare means of IMC

thickness of the BGA IMC layer ...................................................................... 69 Table 21: Results of ANOVA test for equal variance and compare means of IMC

thickness of the QFP IMC layer ........................................................................ 70

Table 22: IMC type ........................................................................................................... 71 Table 23: Sample of temperature cycling requirements Table 4-1 in IPC-9701A

6 .......... 81

Table 24: Worst case use environments of SMT6 ............................................................. 81

Table 25: Build matrix for ATC testing ............................................................................ 89 Table 26: Monitored components for ATC testing ........................................................... 90

Table 27: Test matrix for 0C to 100C ATC ................................................................... 91

Table 28: Test matrix for -50-125C ATC ....................................................................... 92

Table 29: Summary of QFP failures after 0C to 100C ATC ......................................... 95

Table 30: Summary of BGA failures after 0C to 100C ATC ........................................ 99

Table 31: Summary of QFP failures after -55C to 125C ATC .................................... 102

Table 32: Summary of BGA failures after -55C to 125C ATC ................................... 105

Table 33: Results of Levine-Test to compare the variance () of IMC measurement and

2-sided t-Test to compare the means (µ) of IMC measurements .................... 115

Table 34: Results of Levine-Test to compare the variance () of Cu3Sn measurement and

2-sided t-Test to compare the means (µ) of Cu3Sn measurements ................. 115 Table 35: Results of ANOVA test to equal variance and compare means ..................... 116 Table 36: CTE values for common materials in solder joints ........................................ 131 Table 37: JESD22A121.01 test conditions

4 .................................................................... 135

Table 38: Alloys screened for whisker growth ............................................................... 136

viii

Table 39: Summary of whisker growth after 1610 cycles thermal shock....................... 140


thickness of the QFP IMC layer after 1610 cycles thermal shock. ................. 149 Table 41: Build matrix for drop shock testing ................................................................ 158

Table 42: Monitored components for drop testing ......................................................... 160 Table 43: Drop test results .............................................................................................. 166 Table 44: Failures of BGA on High Tg boards ............................................................... 171 Table 45: Failures of BGA on Normal Tg boards .......................................................... 174 Table 46: Solder paste test matrix ................................................................................... 183

Table 47: ReMap M2 test matrix .................................................................................... 186 Table 48: ReMap M3 test matrix .................................................................................... 189

ix

List of Figures

Figure 1: Schematic of a BGA solder joint ......................................................................... 1 Figure 2: Image of a QFP solder joint

3 ............................................................................... 2

Figure 3: Pad crater defect seen in a) cross section – indicated by arrows, and b) viewed

from above .......................................................................................................... 4 Figure 4: Ag3Sn platelets seen in a) cross sectioned solder joint and b) SEM image of

solder joint after removal of Sn phase .............................................................. 5 Figure 5: Sn Whisker growing from SAC305 solder ......................................................... 6 Figure 6: Sn-Cu phase diagram........................................................................................... 7 Figure 7: a) SAC ternary phase diagram and b) Sn-rich corner of SAC ternary phase

diagram ............................................................................................................. 11

Figure 8: Head in pillow defect ........................................................................................ 12 Figure 9: a) gull wing leaded solder joint, b) BGA solder joint ....................................... 13

Figure 10: Solder joint formed with leadless component ................................................. 15

Figure 11: SnBi phase diagram ......................................................................................... 16 Figure 12: a) SnAgBi ternary phase diagram and b) Sn-rich corner of SnAgBi phase

diagram ............................................................................................................. 17 Figure 13: a) SnCuBi ternary phase diagram and b) Sn-rich corner of SnCuBi phase

diagram ............................................................................................................. 18

Figure 14: Example of ternary phase diagram31

............................................................... 19 Figure 16: DSC scan of Sn20%Bi

10, ................................................................................. 22

Figure 17: Microstructure of solder joint formed with SAC305 solder ball and a no-Ag

solder paste alloy6 ............................................................................................. 25

Figure 18: Honeywell test vehicle, medium complexity board ........................................ 26

Figure 19: Example of wetting on OSP, QFP240 ............................................................. 28

Figure 22: New approach proposing to use solder with no or low-Ag content ................ 35 Figure 23: Typical reflow profile for SAC305 ................................................................. 42 Figure 24: Copper dissolution of BGA solder joint after a) 1 replacement b) 3

replacements and c) 5 replacements ................................................................. 43 Figure 25: SnAgCu phase diagram with SAC305 equilibrium solidification path .......... 44

Figure 26: A 3D phase diagram of Sn rich portion of Sn-Ag-Cu ternary system ............ 46 Figure 27: SAC solder joint viewed a) in cross section and b) after selective

electrochemical etching to remove Sn phase .................................................. 48

Figure 28: Solder joint between Sn-based solder and Cu a) rod shaped Cu6Sn5 () ...... 50

Figure 29: SnPb solder jointed to Cu substrate a) showing Kirkendall voids b) showing

Pb phase pooling at IMC/solder interface (cross section by Zohreh Bagheri) . 51

Figure 30: Typical morphologies of Cu6Sn5 grains formed on single crystal ................. 52 Figure 31: IMC formed between SAC solder and Ni(P)Au

23 ........................................... 54

Figure 32: Test vehicle...................................................................................................... 55 Figure 33: Area for EDX compositional analysis of a) bulk BGA solder joint, ............... 56 Figure 34: Optical images QFP a) SAC305 b) Senju M42 c) Sunrise and d) Sunflower . 59

Figure 35: Bi present in a) eutectic colonies in Sunrise and b) throughout bulk solder

(eutectic colonies and Sn dendrite arms) of Sunflower .................................. 59

x

Figure 36: QFP solder joint as seen a) optically b) SE SEM and c) BSE SEM ............... 60

Figure 38: SEM image of Sunflower BGA solder joint ................................................... 63 Figure 39: EDX mapping of intermetallic in Sunflower showing a) image ..................... 64 Figure 42: Polarized light images BGA a) SAC305 b) Senju M42 .................................. 67

Figure 44: IMC measurements of QFP solder joints after reflow (U2) ............................ 70 Figure 45: IMC bond layers formed on a) QFP components and b) BGA components ... 71 Figure 46: IMC formed on BGA at component side ........................................................ 72 Figure 47: Typical morphology of a) Cu6Sn5 IMC as formed on a QFP solder joint with

Sunflower b) and c) (Cu,Ni)6Sn5 and Ni23Cu33Sn44 IMCs respectively both

formed on a BGA with SAC305 solder paste ................................................... 72 Figure 49: Hysteresis loop for thermal cycle .................................................................... 80

Figure 50: SnPb solder a) before testing and b) after 3000 cycles -55C to 125C ......... 83

Figure 51: SAC305 solder a) before testing and b) after 3000 cycles -55C to 125C12

. 84

Figure 52: SAC305 after ATC shown with a)polarized light and b)EBSD mapping 13

... 85 Figure 53: Test vehicle with monitored components ........................................................ 89

Figure 54: Card set up in thermal cycling chamber .......................................................... 90

Figure 55: Chamber profile for 0C to 100C thermal cycling ........................................ 91

Figure 57: Weibull plots of SAC305 QFP solder joints after 6010 Cycles 0 to 100C

comparing High Tg to Normal Tg boards .......................................................... 96 Figure 58: Probability Density Function for SAC305 QFP solder joints after 6010 Cycles

0 to 100C comparing High Tg to Normal Tg boards ...................................... 96 Figure 59: QFP fracture of SAC305 on High Tg board a) optically and b) cross section 97

Figure 60: Fracture surface of QFP solder joint with SAC305 on High Tg board after

6010 cycles ........................................................................................................ 98 Figure 61: Fracture initiation in QFP a) Sunflower b) Senju M42 after 6010 cycles ....... 99

Figure 62: Weibull plots of BGA solder joints on Normal Tg boards after 6010 Cycles 0

to 100C cycles comparing Four Alloys ......................................................... 100

Figure 63: BGA failures after 0 to 100C ATC a) failure in board material by via plating

crack and b) partial failure through bulk solder (SAC305) near component side

IMC ................................................................................................................. 101

Figure 64: Weibull plots of QFP failures on High Tg boards after 1000 cycles -55 -125C

103

Figure 65: Probability Density Function for QFP failures on High Tg boards after 1000

cycles -55 -125C ............................................................................................ 103 Figure 66: Weibull plots of QFP failures on Normal Tg boards after 1000 Cycles -55-

125C .............................................................................................................. 104 Figure 67: QFP176 fractures in a) Sunflower and b) Senju M42 after 1000 cycles Harsh

testing .............................................................................................................. 105

Figure 68: BGA failures after -55 to 125C ATC failures in board material by via plating

crack in a) boards built with Sunrise and b) boards built with Sunflower ...... 106 Figure 70: Weibull plots of BGA solder joints after 1000 Cycles -55 to 125C cycles

comparing two board materials ....................................................................... 107 Figure 71: Weibull plot for QFP solder joints on Normal Tg boards comparing two test

conditions ........................................................................................................ 108

xi

Figure 72: Probability Density Function for QFP failures on Normal Tg boards

comparing two test conditions ........................................................................ 108 Figure 73: SAC305 at Time 0 a) optically and b) SEM image and after b) 3148 cycles

and ................................................................................................................... 109

Figure 74: Senju M42 at 1000x after a) 3148 cycles and b) 6010 cycles ....................... 110

Figure 75: Sunrise after 6010 cycles 0-100C at a) 500x and b) 1000x ......................... 110

Figure 76: Sunflower at a) Time 0 and b) after 6010 cycles 0-100C at 500x ............... 111 Figure 77: IMC layers formed on QFP176 solder joints between the board side Cu layer

and solder paste a) sunrise and b) SAC305 after 438 cycles of Harsh thermal

cycling. Location 1 shows the Cu3Sn layer, location 2 shows the Cu6Sn5 layer

112 Figure 78: IMC growth at the board side of QFP during thermal cycling ...................... 114 Figure 80: Interval plot of IMC thickness at the board side of QFP after ATC ............. 117

Figure 81: Interval plot of Cu3Sn thickness at board side of QFP after ATC ................ 117

Figure 83: Main effects plot of IMC thickness at the board side QFP during 0-100C . 119

Figure 85: Main effects plot of Cu3Sn thickness board side of QFP during 0-100C ... 120 Figure 86: Interaction plot of IMC thickness at the board side of QFP after ATC ........ 121 Figure 88: Pb "cap" on whisker from SnPb component finish

2 ...................................... 127

Figure 89: a) Schematic for a typical solder joint of a leaded component using SnPb

solder, b) cross section showing solder joint formed with Pb-free solder ...... 128

Figure 90: Cyclic Dynamic Recrystallization resulting in whisker formation ............... 129 Figure 91: Source of compressive stress contributing to whisker growth ...................... 130 Figure 92: Ag3Sn oxide zone with whisker

6 .................................................................. 132

Figure 93: Whisker morphology a) long, thin whiskers and b) short, kinked whiskers16

133

Figure 94: Cross sections of plating surface made with SEM FIB of a) Sn and b) SnPb19

133

Figure 95: Samples in HTHH chamber .......................................................................... 136 Figure 96: Two stage, air to air, chamber for thermal shock testing .............................. 137

Figure 97: Thermal shock temperature profile ............................................................... 138 Figure 98: Schematic showing locations on lead where whiskers formed ..................... 140

Figure 99: Two adjacent leads with Sunflower solder paste after 1610 thermal shocks 141 Figure 100: Whisker growth on SAC305 after 1610 thermal shocks. a) and b) whisker

growth in location 4 c) massive deformation and d) whisker growth in location1

142 Figure 101: Whisker growth on Senju M42 after 1610 thermal shocks. a) and b) location

4 with short, thick whisker surrounded by many, very short whisker nucleation

sites c) and d) short, thick whiskers growing at site of contamination in

location4 .......................................................................................................... 143

Figure 102: Whisker growth on Sunrise after 1610 thermal shocks. a) and b) location 2

with short, thick whisker c) location 4 and d) location 1 with many, very short

whisker nucleation sites short ......................................................................... 144 Figure 103: Whisker growth on Sunflower after 1610 thermal shocks. a) and b) location 1

with short, thick whisker c) some, very short whisker nucleation sites at

location 4 and d) longest whisker observed at location1 ................................. 145

xii

Figure 104: Whisker growing from Sunflower at grain boundary of recrystallized grains

146 Figure 105: Hillock growing from Sunrise after thermal shock ..................................... 146 Figure 106: Senju M42 bulk recrystallization ................................................................ 147

Figure 107: IMC measurements of QFP solder joints after 1610 cycles thermal shock

(U1) ................................................................................................................. 148 Figure 108: Interval plot of IMC thickness at the lead side of QFP after thermal shock 149 Figure 109: IMC layer at lead with a) SAC305 and b) Sunrise after 1610 cycles of

thermal shock .................................................................................................. 150

Figure 111: Bi content at various locations of a Sunrise solder joint ............................. 151 Figure 112: Cross section of Sunflower showing Bi accumulating at grain boundaries 152 Figure 113: Test vehicle with monitored components .................................................... 159 Figure 114: Solder mask defined vs. non-solder mask defined ...................................... 160

Figure 115: Example of accelerometer secured with RTV silicone ............................... 162 Figure 116: Test set up .................................................................................................... 162

Figure 117: Target pulse shock defined by JESD22-B110 service condition B............. 163 Figure 118: Sample of pulse shock achieved during test ................................................ 164

Figure 119: Individual value plot of drops to fail, BGA (SAC305 + alloy) (U205) ...... 167 Figure 120: Individual value plot of drops to fail, QFP (U2) ......................................... 167 Figure 121: Failure modes of solder joint as defined by IPC/JEDEC-9702 ................... 169

Figure 122: BGA failure by pad cratering, Sunflower on High Tg boards ..................... 170 Figure 123: D&P mapping of Senju M42 on High Tg board, with images of ................ 171

Figure 124: Schematic of Quad-Flat-Package (QFP) ..................................................... 172 Figure 125: QFP failures in Sunflower on High Tg boards ............................................ 172 Figure 126: D&P mapping of Sunrise on Normal Tg board, with images of ................. 173

Figure 127: BGA failure in Sunflower on Normal Tg boards ......................................... 174

Figure 128: QFP failures in Sunflower on Normal Tg boards ........................................ 175 Figure 129: QFP failure in Sunrise on Normal Tg board ................................................ 175 Figure 130: Orchid at a) Time 0 and b) after 3000 cycles ATC ..................................... 187

Figure 131: Sunrise solder joint after 1000 hrs HTHH followed by 2 years ambient

storage ............................................................................................................. 188

xiii

List of Acronyms

AF: Accelerating Factor

Ag: Silver

ALT: Accelerated Life Testing

ANOVA: analysis of variance

ATC: Accelerated Thermal Cycling

ATHH: Ambient Temperature High Humidity

Au: Gold

BGA: Ball Grid Array

Bi: Bismuth

BSE: Backscatter Electron

CBGA: Ceramic Ball Grid Array

Cd: Cadmium

CTE: Coefficient of Thermal Expantion

Cu: Copper

D&P: Dye and Pry

DI: De-Ionized

DRX: Dynamic Recrystallization

DSC: Differential Scanning Calorimeter

EDX: Energy Dispersive X-Ray Spectroscopy

ENEPIG: Electroless Nickel Electroless Palladium Immersion Gold

ENIG: Electronless Nickel Immersion Gold

FA: Failure Analysis

FR4: flame retardant level

H0: null hypothesis

Ha: alternate hypothesis

HiP: Head in Pillow

HTHH: High Temperature High Humidity

ICT: In Circuit Testing

IMC: Intermetallic Compound

In: Indium

iNEMI: International Electronics Manufacturing Initiative

JEDEC: Joint Electron Device Engineering Council

LQFP: Low Profile Quad Flat Pack

MLF: Micro Lead Frame

MTTF: Mean Time To Failure

Ni: Nickel

NTC: Number of Thermal Cycles

OSP: Organic Solderability Preservatives

Pb: Lead

Pb-free: Lead free

PBGA: Plastic Ball Grid Array

PDF: probability density function

PWB: Printed Wire Board

xiv

QFP: Quad Flat Package

RoHS: Restriction of Hazardous Substances

SAC: Tin-Silver-Copper (SnAgCu)

SE: Secondary Electron

SEM: Scanning Electron Microscope

SiC: Silicon Carbide

SMT: Surface Mount Technology

Sn: Tin

SnPb: Tin Lead (usually eutectic Tin Lead)

Tg: Glass Transition Temperature

Th: Homologous Temperature

Tm: melting temperature

Vacc: Acceleration Voltage

Zn: Zinc

1

Chapter 1 Background

1 Introduction

A solder joint is formed when two metal conductors with high melting temperatures,

usually copper (Cu) which melts at 1085C, are joined with another metal having a low

melting temperature. Tin (Sn) comprises the majority of materials used as solder because

of its low melting temperature (231.9°C), and its unique ability to form intermetallics

with many different joining substrates.1 In the case of electronic surface mount

technology (SMT), which is the focus of this study, these bonds act as electrical,

mechanical and thermal bonds. The solder joint is comprised of two interfaces between

the copper and solder as well as the bulk solder, which is filler material between the two

interfaces. The interfaces are chemical bonds between the Sn from the solder and the Cu

from the connecting conductors.2 In some cases, an intermediary nickel (Ni) layer will

also be present. The intermetallic compound (IMC) layer, which forms during the reflow

and solidification processes, tends to be more brittle than either the Cu or the bulk solder.

Figure 1 is a schematic of a typical ball grid array (BGA) type of component solder joint

and Figure 2 is an image of a solder joint formed using a quad flat package (QFP) type of

leaded component with gull winged shaped leads. Both are attached to the printed wire

board (PWB) by the formation of a solder joint. The solder joints formed using these two

types of components represent two of the main ways in which solder is used to form

electrical contacts.

Figure 1: Schematic of a BGA solder joint3

2

Figure 2: Image of a QFP solder joint3

2 Need for New Low Melt Solder

2.1 Legislative Changes

In 2006 The Restriction of Hazardous Substances Directive or RoHS (2002/95/EC) took

effect for electronic products put onto the market across the European Union.4 Each of

the six substances restricted within this legislation (with some exemptions) pose some

degree of challenge to the electronic manufacturing industry. However, it is the ban of

lead (Pb) which presents the greatest disruption to traditional manufacturing processes.

Some 50 years of manufacturing knowledge based on eutectic tin-lead (SnPb) solder

requires revision, update and change. Due to the global nature of the electronic

manufacturing sector, this change has impacted all electronic manufacturing whether the

products are destined for European markets or elsewhere.

Furthermore, the European Union revised the legislation in 2011, as part of the RoHS

recast,5 and expressed an ongoing commitment to review legislation regularly, in

particular material exemptions. Any replacement to SnPb solder therefore needs to

consider the toxicity and downstream environmental impacts of the new materials.

3

2.2 SAC305 and Current Issues with Existing Alloys

The search for a replacement alloy for electronic assemblies has lead to the current

industry standard of tin-silver-copper alloys, (Sn-Ag-Cu or SAC); the most widely used

are SAC305 and SAC405, both contain 0.5wt% Cu and 3wt% and 4wt% silver (Ag)

respectively. Although these Sn-based alloys are in wide use, there remain a number of

technical and environmental limitations, which necessitate the continued search for a

more ideal replacement to SnPb. Some of the key issues with SAC305 and SAC405 are

summarized below.6

Both SAC305 and SAC405 are non-eutectic alloys with melting ranges of 217-220C and

217-221C respectively; considerably higher than the 183C melting point of eutectic

SnPb. Process temperatures for SAC alloys currently run at 240C, which is higher than

the typical process temperatures of SnPb (220C). This increase in process temperature

results in higher energy usage of reflow ovens, and consequently increased CO2

emissions, negating some of the environmental benefits of a less toxic substance.

The process temperature employed for manufacturing with SAC alloys is much closer to

the melting range, narrowing the process window and thus requiring greater process

control. While SnPb is routinely run at approximately 40C above the solder melting

temperature, such a superheat temperature would prove damaging to the overall product

when applied to products using SAC alloys. Both the PWB laminate material and some

of the more temperature sensitive components would not be able to withstand these

higher temperatures. Kelly et al.7 showed that in standard SAC reflow processes, the

component body temperature reached as high as 244.5C. It is expected that smaller

components may reach temperatures as high as 250C. Rework processes, or any wave

soldering, may require even higher temperatures, up to 260C, for compatibility with the

current Pb-free processes.8 Component suppliers have indicated that moisture sensitivity

of components may change in response to these high temperature requirements and may

necessitate preconditioning before reflow. All additional processes increase production

cost and potentially compromise the overall component reliability.

4

Moreover, the laminate materials typically used in conjunction with SnPb solders have

been replaced by new materials which have a higher glass transition temperature (Tg) in

order to withstand the higher processing temperatures needed for SAC alloys. The

development of these new materials involved changing the epoxy resin and curing agents

for additional crosslinking, as well as the inclusion of ceramic particle fillers.9 The

unintended consequence of these changes has been the introduction of a new failure

mode known as pad cratering. The harder and stiffer laminate materials fail as they

separate from the conductive copper traces within the PWB (Figure 3). These laminates

are also more susceptible to warpage, delamination and may require additional baking

processes.

Figure 3: Pad crater defect seen in a) cross section – indicated by arrows, and b)

viewed from above

This failure mode is dangerous as it may not be readily identifiable during in- circuit-

testing (ICT) within the factory, but may lead to reliability concerns in the field as the

fracture propagates, potentially severing the copper trace and causing eventual electrical

failure. This is primarily a concern for products required to withstand damage by drop

shock, such as consumer electronics (e.g. mobile phones), which are prone to being

repeatedly dropped over the course of useful life. It would therefore be preferable for any

new alloy to be compatible with older laminate materials similar to those used with

traditional SnPb solders, in which this failure mode was not prevalent.

5

The mechanical properties of SAC solder alloys also have disadvantages. In addition to

requiring a laminate material with a higher elastic modulus (E), the alloy itself is stiffer

than SnPb.10

. SnPb has an E value of 40GPa at 25C and SAC305 has a value 50GPa at

25C.11

Finally, there is potential for the formation of undesirable intermetallic

compounds (IMC), i.e. Ag3Sn (Figure 4). This intermetallic may in some cases be

present as a large, sharp platelet which can potentially behave as a stress raiser or provide

a path along which cracks may more easily propagate, thereby negatively impacting the

materials’ resistance to drop shock.6

In addition to the aforementioned thermo-mechanical issues, the high cost of Ag makes

its inclusion as an alloying element undesirable for the cost-sensitive market of consumer

electronics.

Figure 4: Ag3Sn platelets seen in a) cross sectioned solder joint and b) SEM image of

solder joint after removal of Sn phase12

Finally, Sn whiskers discussed in detail in Chapter Chapter 4, present a reliability

concern for all Sn-based solders. Although the presence of Pb in high concentration does

not eliminate the potential for growth of Sn whiskers completely, Pb does mitigate the

formation of long whiskers, and therefore is not considered a reliability risk. The move

to high Sn solder compositions coincides with the further miniaturization of components,

solder joints and solder joint spacing. These factors all increase the risk of whisker

growth to a length at which bridging and hence short-circuiting of solder joints would be

6

a concern. While this risk presents a reliability concern for all products using Pb-free

solders, it is of particular concern to the aerospace and defense sectors, for which these

reliability concerns present an unacceptable level of risk.10

Figure 5 shows a Sn whisker

growing from SAC305 solder on the surface of a solder joint.

Figure 5: Sn Whisker growing from SAC305 solder

In spite of these concerns, at the time of RoHS implementation in 2006, near-eutectic

SAC alloys provided the best compromise in a supply chain where Pb containing

components used in printed circuit boards were still common.

2.3 Search for Replacement Solder

2.3.1 Binary Alloys

Prior to implementing SAC as an alternative to SnPb, an attempt to find an optimal

replacement began with the exploration of binary alloy options. A number of binary

alloys were initially considered (ideally near the eutectic point) but all had associated

issues, which disqualified them as viable alternatives.13,14

Tin (Sn) was selected as one of the two binary elements in all cases due to its low

toxicity, relatively low cost, resistance to oxidation and contribution to the melting point

or range of an alloy. Much of this analysis was performed by reviewing the equilibrium

phase diagrams of various Sn-based alloys. Information pertaining to the temperature at

7

the eutectic point, as well as IMCs that may form, was weighed against a number of

factors associated with the individual alloying elements.

Figure 6: Sn-Cu phase diagram15

SnCu has a eutectic point at 0.7% Cu close to the melting temperature of pure tin at

approximately 227C, (Figure 6) which is still more than 40C higher than that the SnPb

eutectic temperature.16

While this alloy presents some advantages, the melting

temperature is still too high to be considered a viable binary compound alternative.

Regardless, the SnCu alloy must be considered in the development of a replacement alloy

due to the interaction of solder with a copper pad on the PWB. In some cases, there will

be a barrier between the copper pad and the Sn-based solder, for example a nickel (Ni)

barrier layer. In other cases, the Sn-based solder will be exposed directly to a solid

copper substrate, which would represent the “worst case scenario” in terms of copper

dissolution. Whichever alloy is selected to replace SnPb, it is likely that the Sn-Cu

reaction will play a large role. This will be discussed further in Chapter Chapter 2.

8

Other binary alloys formed with Sn were also considered and dismissed for several

reasons. A summary of Sn-based eutectic alloys, their melting temperatures and their

eutectic compositions is provided (Table 1).

Table 1: Sn-based eutectic alloys17

System Eutectic Temperature (C) Eutectic Composition (wt%)

Sn-Cu 227 0.7

Sn-Ag 221 3.5

Sn-Au 217 10.0

Sn-Zn 198.5 9.0

Sn-Pb 183 38.1

Sn-Cd 177 32.2518

Sn-Bi 139 57.0

Sn-In 120 51.0

SnIn, SnAg and SnAu all present very expensive options. The alloying elements are

comparatively scarce and therefore do not provide attractive alternatives to SnPb.

Further, the eutectic composition of SnAg and SnAu, as well as that of SnCu is

predominately Sn. These systems will therefore form intermetallic compounds within the

Sn matrix, introducing inhomogeneity into the microstructure.17

Additionally, Ag3Sn

precipitates, which are primarily present as a network of small, rod like structures, may

also form as an undesirable platelet structure as already described above and illustrated in

Figure 4.

SnZn and SnCd both have melting temperatures close to that of SnPb, however both have

been disqualified as viable options. SnZn readily oxidizes;10

this alloy would require

processing in an inert gas environment and the addition of an excessive amount of flux.

Since the oxide forms so easily, it would also potentially form a very weak solder joint.

The main issue with SnCd is the toxicity of Cd; any replacement to SnPb should not

present the same environmental concerns as the original alloy.

9

The binary SnBi eutectic alloy was initially rejected as a viable replacement to SnPb

because of the possibility of forming a dangerously low melting intermetallic compound

when combined with Pb. Moon et al.19

showed that even a 0.1% Pb contamination

content was sufficient to reduce the melting temperature by between 38 and 55C.

During the initial transition to lead free solders, it was anticipated that this contamination

could be introduced when using SnPb tinned leaded parts within a lead-free assembly.

The potential for this contamination within the supply chain was deemed to be too great

and therefore all bismuth (Bi) containing alloys were avoided. As the legislation has now

been in place for almost a decade, this contamination has been deemed to be less likely to

occur and therefore Bi is being reconsidered a suitable alloying element.20

Research on

eutectic SnBi alloy continues. It has also been has been adopted in high volume

production for special applications; flat screen televisions are built using a combination

of SAC and eutectic SnBi. However, there remains concern due to the brittle nature of the

eutectic and its resulting poor performance in drop shock testing;21

this study will instead

consider Bi as an alloying element within a predominately Sn-based alloy system.

2.3.2 Ternary and higher order alloy options

Once it was determined that binary eutectic Sn-based alloys would not meet the needs of

a SnPb replacement, ternary alloys were explored. The International Electronics

Manufacturing Initiative (iNEMI) proposed the following criteria for the selection of a

replacement alloy:

“Have melting point as close to Sn-Pb eutectic as possible

Be eutectic or very close to eutectic

Contain no more than three elements (ternary composition)

Avoid using existing patents, if possible (for ease of implementation)

Have the potential for reliability equal to or better than SnPb eutectic.”22

Further criteria were added for narrowing down a suitable ternary alloy. These include22

:

Liquidus temperature as close as possible to 183C

Solidus temperature as close as possible to liquidus temperature (small pasty range)

10

Solidus temperature significantly higher than the solder joints maximum operating

temperature

SnAgCu, SnBiCu, and SnAgBi were all considered. Based on the possibility of forming

very low melting temperature eutectics in the presence of Pb (BiPb eutectic at 125.9C,

SnPbBi ternary eutectic at 98C), Bi was ruled out during the initial transition period and

SnAgCu alloys became standard. However, now that the supply chain has been largely

purged of SnPb components, Bi containing alloys are being reconsidered. Bi-rich alloys

continue to be considered for some specialty applications, for example Sn-88%Bi and Sn-

59%Bi-1.2%Ag. However, as they have a melting temperature closer to 138C, they fail

the third criteria listed above for applications which may be exposed to harsh operating

environments of up to 125C.

The SnAgCu eutectic system is of the composition Sn-3.5Ag-0.9Cu by wt% with a

eutectic melting temperature of 217.4±0.8C. The system forms a ternary eutectic made

up of faceted Cu6Sn5 and non-faceted Ag3Sn internetallics within a Sn matrix. The

melting temperature of this eutectic is 10C lower than that of eutectic SnCu. Near

eutectic compounds of SnAgCu are currently the preferred solders in use today,

specifically SAC305, SAC405 and to a lesser extent SAC105. Where cost is a driving

force (i.e. consumer products) the preference is for a lower quantity of Ag. At the other

end of the spectrum, Ag is limited due to the potential formation of undesirable Ag3Sn

platelets.

Figure 7 shows the phase diagram of SAC alloys and specifically the Sn-rich region,

which highlights the formation of a Sn - Sn+Ag3Sn-Sn - Cu6Sn5 ternary compound.

11

Figure 7: a) SAC ternary phase diagram and b) Sn-rich corner of SAC ternary

phase diagram23

SAC105 has been considered as an alternative to SAC305 or SAC405 as it reduces the

amount of expensive Ag, and also reducing the possible formation of the dangerous

Ag3Sn platelets. However, the reduction of Ag also coincides with an increase in

liquidus temperature to 227C.6 The corresponding increase in required reflow, rework

and other processing temperatures offsets any potential benefit gained by this slightly less

stiff, higher fracture toughness alloy. The introduction of this alloy into the ball grid

array (BGA) supply chain, where it was combined with a paste alloy of SAC305 or SnPb

(in the case of mixed metallurgy builds early in the transition to Pb-Free) lead to issues of

incomplete melting of the solder ball. This incomplete melting resulted in a Head-in-

Pillow (HiP) defect (Figure 8), since the manufacturing reflow temperatures were

generally set to accommodate the SAC305 paste.

b. a.

12

Figure 8: Head in pillow defect24

2.4 Celestica’s Low Melt Solder Research Program

Celestica, in partnership with the University of Toronto, set up a low melt research

program in 2009 with the main goal of introducing a solder alloy which would allow for

the use of conventional flame-retardant level 4 (FR4) laminate printed wire board (PWB)

material with SnPb solder. This combination would require a process temperature of

approximately 10C lower than SAC305 and potentially reduce failures via pad cratering.

This new solder paste would need to be compatible with both leaded and discrete

components - where the solder paste makes up the majority of the solder joint, and with

ball grid array components (BGA) where the paste is mixed with industry standard

SAC305 or SAC105 to form a solder joint which has the composition made up of

approximately 25% by volume paste, and approximately 75% by volume BGA alloy10

(Figure 9).

13

Figure 9: a) gull wing leaded solder joint, b) BGA solder joint

Celestica’s program was divided into a number of phases as follows:

Alloy Selection

Manufacturing Feasibility

Screening Experiments:

o Aerospace and Defense Sector

o Consumer Sector

o Telecommunications Sector

Reliability Testing

This work explored the alloys selected for both the Consumer and Telecommunications

sector portions of the screening experiments. The Aerospace and Defense sector program

is described here in order to provide context for the current project. The remainder of

this chapter summarizes the prior work and provides context for the work in this thesis.

2.4.1 Phase 1: Alloy Selection

The initial work, described in Snugovsky et al.6, outlines the process by which various

alloys were selected for further study. This alloy selection phase of the project was

divided further into the following stages:

14

Requirement Formulation

Literature Search and Phase Diagram Analysis

Preliminary Alloy Selection

Metallurgical Analysis: Differential Scanning Calorimetry (DSC) and Microstructural

Evaluation of:

o Alloy

o Alloy plus Cu – simulate solder on board with leaded or leadless

component

o Alloy plus Cu plus SAC305/SAC105 – simulate solder on board with

BGA solder ball

2.4.1.1 Requirement Formulation

The following requirements were established:

220-222C process temperature, at least 10C lower than SAC305

Approximately 10-15C pasty range

Compatible with SAC305/105 solder balls as well as with leaded and discrete

components

Low Ag content6

Equal or better thermo-mechanical properties than SAC305.

As described in section 2.2, an ideal replacement to SAC solders needs to be compatible

with temperature sensitive components. It should also be compatible with laminate PWB

material used with SnPb solder, which is less expensive than newer, high temperature

laminate material, requires less baking and processing prior to build and reduced adverse

mechanical properties associated with these new materials; pad cratering, de-lamination

and warpage.

The pasty range is also important as a narrow range ensures good wettability due to less

oxide formation and reduces the possibility of forming HiP defects if the paste does fully

15

melt and mix with the BGA solder ball. HiP may also result from excessive warpage in

the laminate material. A small pasty range also prevents composition gradients and

unexpected concentrations driven by thermal gradients in the cooling cycle.

Finally, any new low melt solder should form an interconnection which exhibits equal or

better properties in mechanical and thermo-mechanical testing than SAC305 and

comparable to SnPb solder. The interconnection includes the solder paste, the board

material (copper pad and any surface finish material), and either the BGA solder ball, the

lead material of a leaded component (Figure 9), or a terminal finish of a leadless

component (e.g. a capacitor) (Figure 10).

Figure 10: Solder joint formed with leadless component25

The reduction of Ag will reduce the cost of the alloy and also may improve mechanical

properties, particularly in drop shock testing. It should also be noted that the ideal Ag

content for drop shock performance may not be optimal in thermal cycling.26

It is

therefore necessary to balance these requirements.

2.4.1.2 Literature Search and Phase Diagram Analysis6

Bi containing alloys are the focus of this study. It has been shown that the addition of Bi

will improve thermal and thermo-mechanical properties over conventional SAC

alloys.27,28

These benefits can be attributed to solid solution and precipitate strengthening

which will be described further in Chapter Chapter 2. Additionally, Bi may reduce the

propensity for whisker growth in a Sn rich alloy.6

The Bi addition was intended to

depress the melting temperature but not to enter the very low melting temperature regime

around the SnBi eutectic (Figure 11). The Bi content is tuned in conjunction with altering

16

the amount of Ag and/or Cu in order to maximize the temperature depressing

characteristic of the Bi, without introducing the very low melting temperature (~137-

138C) ternary eutectics (Figure 12 and Figure 13).

Figure 11: SnBi phase diagram29

Since alloys of higher order than ternary are difficult to illustrate using phase diagrams,

the combined knowledge of the existing ternary and binary eutectic temperatures and

compositions can be compared to determine an ideal range for the proposed alloys. Since

a melting temperature of approximately 138C is well below the outlined criteria, alloys

which do not have any liquid component around this temperature are viable candidates. A

slice along the z-axis, which represents temperature, of a 3D ternary phase diagram

(Figure 14) can be used to explore the eutectic region of a particular alloy system.

The solubility limit of Bi in Sn is 21%, however this is reduced with the addition of

other alloying elements, namely Ag and/or Cu. An examination of the Sn-Ag-Bi, Sn-Cu-

Bi and Sn-Ag-Cu systems is required.

17

Figure 12: a) SnAgBi ternary phase diagram and b) Sn-rich corner of SnAgBi phase

diagram30

From Figure 12 and Equations 1-1 and 1-2, a ternary eutectic (Point ‘E’) of Sn-Ag-Bi

exists at 137.1C, at a composition of Sn-0.68%Ag-43.47%Bi, and a binary eutectic

(Point ‘A’) at 220.3C with a composition of Sn-3.73%Ag is seen.

L -> Ag3Sn + (Bi) + (Sn) 2-1

L -> Ag3Sn + (Sn) 2-2

Upon cooling along the binary eutectic line from point A, highlighted in blue in Figure

12, both Sn and Ag3Sn are formed simultaneously. At point E (the ternary eutectic point

defined in Equation 1-1) all remaining liquid forms simultaneously into Sn, Ag3Sn and

Bi.31

Compositions along this line were explored in the development of proposed new

alloys.

a. b.

18

Figure 13: a) SnCuBi ternary phase diagram and b) Sn-rich corner of SnCuBi phase

diagram 32

Similarly, from Figure 13 and Equation 1-3, we can see that a ternary eutectic exists at

138.8C at a composition of Sn-0.01%Cu-43.09%Bi and a binary eutectic at 226.8C

with a composition of Sn-0.89%Cu.

L -> Cu6Sn5 + (Sn) + (Bi) 2-3

L -> Cu6Sn5 + (Sn) 2-4

Exploring the binary eutectic line following cooling towards the ternary eutectic from

point A in Figure 13 to point E for the SnCuBi system gives a starting point for

determining ideal alloys.

Figure 15 is an isothermal projection of the SnBiAg phase diagram at 138.4C. The area

highlighted in yellow represents an area where ternary eutectic is present at this

temperature. Compositions from within these triangles should be avoided in order to

avoid melting in the low temperature range. Additionally, the ternary eutectic of the Sn-

Ag-Bi system should be avoided because Bi is present in the eutectic as a primary

precipitate. These primary particles may be large. It is preferable to have small Bi

a. b.

19

particles, the type which will likely precipitate out of solid solution after initial

solidification.

Finally, it should also be noted that current production processes such as reflow ovens

and hand solder rework, provide rapid cooling environments and do not allow for

equilibrium solidification. Other possible effects of cooling, such as Bi segregation,

should be accounted for in the development of any new alloy.

Figure 14: Example of ternary phase diagram31

20

Figure 15: SnAgBi ternary phase diagram, isothermal section33

2.4.1.3 Preliminary Alloy Selection6

Based on examination of phase diagrams as described in section 2.4.1.2, 23 Bi-containing

alloys were selected for further evaluation including those listed in Table 2. The first two

alloys in the table, J.Hwang and Senju M42, are currently available in industry. Of these

four industry available alloys, none fully meet the criteria outlined in section 2.4.1.1 but

were included for study because they met most criteria and for purposes of comparison.

21

Table 2: Preliminary alloy selection10,6

Alloy Composition

(wt%)

Published Melting

Temperatures (C)

Experimental Melting

Temperatures (C)

Min Max Pasty

Range Min Max

Pasty

Range

J. Hwang34

Sn3.1%Ag0.5%Cu3.1%Bi 209 212 3 209 216 7

Senju M4235

Sn2%Ag0.75%Cu3%Bi 207 218 11 207 217 10

Paul27,28

Sn3.4%Ag4.8%Bi 201 215 14 206 217 11

Orchid Sn2%Ag7%Bi 191 216 25 190 215 25

Violet Sn2.25%Ag0.5%Cu6%Bi 205 216 11

Sunflower Sn0.7%Cu7%Bi 200 219 19 209 219 10

Cornflower Sn0.7%Cu10%Bi 191 215 24

Sunrise Sn1%Ag0.7%Cu7%Bi 199 215 16

2.4.1.4 Metallurgical Analysis: DSC and Microstructural Evaluation6

Alloys from Table 2 were evaluated using the thermoanalytical technique, differential

scanning calorimetery (DSC). DSC measurements of the alloys were made in two forms:

alloy alone representing the solder paste used in the production of a typical leaded or

leadless solder joint, and the alloy combined with SAC305 in approximately 25%-75%

ratio to represent the solder joint formed when the paste is used with a BGA component

(Figure 9). DSC measures the difference in energy input into a cell containing the sample

within a sample holder vs. a reference cell, usually the sample holder (minus the

sample).36

By plotting the heat flow (J/g*min) vs. temperature (C) it is possible to

determine the onset of a phase transformation by the decrease in heat flow, the

temperature range over which this transition occurs and the total enthalpy of the reaction.

In the case of solder in this experiment, the onset of melting and the pasty range could be

22

determined. Table 2 summarizes both the published values of melting temperature and

range as well as values determined in this study. DSC can also be used to determine if

there are any additional, undesirable phase changes that occur in the alloys. As the

selected alloys were close to the eutectic compositions, only one peak is desirable. Figure

16 is a good example of an alloy with an undesirable hypoeutectic composition. Note that

the saturation point of Bi in Sn is at 21wt% Bi. The eutectic composition is

approximately 57wt% Bi. As the alloy is heated from room temperature, the first peak is

encountered at approximately 139C (point A), which is the solid solubility limit for Bi in

Sn as well as the initiation of melting. This is characterized on the DSC curve by a sharp

peak indicating an invariant phase change. Continued heating of this alloy shows that it

passes through a melting range characterized on the DSC curve by a secondary peak, at

point B (liquidus).37

Figure 16: DSC scan of Sn20%Bi10,38

Table 3 summarizes the DSC analysis performed on the proposed alloys combined with

75% SAC305 to simulate the composition of the solder paste mixed with a SAC305

solder ball of a BGA.

23

Table 3: DSC analysis of proposed alloys with 75% SAC305

Alloy Composition

(wt%)

Experimental Melting

Temperatures (C)

Min Max Pasty Range

J. Hwang39

Sn3.1%Ag0.5%Cu3.1%Bi 216 218 2

Senju M4240

Sn2%Ag0.75%Cu3%Bi 216 218 2

Paul27,28

Sn3.4%Ag4.8%Bi 214 217 3

Orchid Sn2%Ag7%Bi 214 218 4

Violet Sn2.25%Ag0.5%Cu6%Bi 214 217 3

Sunflower Sn0.7%Cu7%Bi 212 215 3

Sunrise Sn1%Ag0.7%Cu7%Bi 211 215 4

Microstructural evaluation was performed by cross sectioning the alloyed samples and

examining the microstructure using a scanning electron microscope (SEM) and energy-

dispersive X-ray spectroscopy (EDX). This evaluation was done using three forms of the

proposed alloys:

Alloy

Alloy reflowed onto a copper foil

Alloy reflowed onto a copper foil with SAC305

The second and third scenarios represent the alloy reflowed onto a copper pad found on a

PWB and a BGA mixed with the alloy onto a copper pad respectively. Using the above

three scenarios, each of the alloys were evaluated based on the following microstructural

factors: intermetallic thickness and morphology, microstructural uniformity and

coarseness of the overall microstructure.

The intermetallic formed between the solder alloy and the copper pad creates both the

mechanical and electrical bond of the solder joint. The formation is discussed further in

Chapter Chapter 2. Both Cu6Sn5 and Cu3Sn are possible intermetallics that form during

solder reflow, and are substantially more brittle than Sn or Cu alone and have facetted

24

morphologies. During a typical reflow of Sn based solder onto a copper foil, it is

typically the Cu6Sn5 IMC layer in a scalloped morphology which forms the interconnect.

A continued solid-state reaction between the Sn in the solder, the Cu6Sn5 layer and the Cu

may persist at elevated temperatures (including those present in service conditions)

causing the Cu6Sn5 layer to grow and possibly resulting in the formation of a Cu3Sn layer

between the Cu6Sn5 and Cu over time. In this study, the initial IMC formation processes

were compared. There should be a sufficient, continuous IMC formed with no

interruptions to ensure a proper bond has formed. Conversely, the IMC layer thickness

should be minimized as it is the most brittle portion of the solder joint and prone to

fracture. Thicknesses will vary based on solder composition, PWB surface finish, reflow

profile (time and temperature) and post reflow exposure. An optimal thickness definition

does not currently exist in industry.

Microstructural uniformity and grain coarseness were also examined. The microstructure

plays a role in the materials ability to resist creep by allowing the induced load to be

uniformly distributed over the specimen thereby reducing strain-induced grain growth

and crack initiation.41

Solder joints can readily reach temperatures of approximately 0.3-

0.5 of the melting temperature (Tm) during normal operation. This temperature range may

result in both the coarsening of the primary grain structure, in this case the Sn matrix

grains, as well as growth of Ag3Sn platelets.42

In this study, the alloy itself, as well as the

alloy mixed with SAC305 in a 1:3 ratio were examined to simulate a solder joint formed

with the proposed alloy and a SAC305 BGA solder ball, as shown in Figure 17.

25

Figure 17: Microstructure of solder joint formed with SAC305 solder ball and a no-

Ag solder paste alloy6

2.4.2 Phase 2: Manufacturing Feasibility Study43

Manufacturing feasibility, or solder paste performance, was tested at Celestica. The

pastes were tested for printability, wettability and performance in the screening process.

In general, the performance of a solder paste is driven largely by the size and morphology

of the solder powder, the composition and volume of flux and the overall rheology of the

paste. These variables are largely alloy independent and therefore not considered in detail

during this study. This study concluded that the paste variables have been sufficiently

addressed and optimized in order to move to further stages of evaluation, which would

better distinguish between alloys.

2.4.3 Phase 3: Screening Experiments

From the alloys described in 2.4.1, it was determined that no one alloy would meet the

needs of the three market segments of interest. It was therefore proposed that a number of

screening experiments be performed in order to down-select the number of alloys which

would undergo full reliability studies. Table 4 summarizes the alloys selected for each of

the screening experiments for different market segments.

26

Table 4: Alloys for screening experiments

Market Segment Alloy Composition

(wt%)

Process

Temperature

(C)

Pasty

Range (C)

Aerospace and

Defense

Paul Sn3.4%Ag4.8%Bi 206 11

Violet Sn2.25%Ag0.5%Cu6%Bi 205 10

Orchid Sn2%Ag7%Bi 190 25

Consumer

Sunflower Sn0.7%Cu7%Bi 226 10

Sunrise Sn1%Ag0.7%Cu7%Bi 222 16

Senju M42 Sn2%Ag0.75%Cu3%Bi 224 10

Telecommunications

Violet Sn2.25%Ag0.5%Cu6%Bi 205 10

Sunflower Sn0.7%Cu7%Bi 226 10

Sunrise Sn1%Ag0.7%Cu7%Bi 222 16

Senju M42 Sn2%Ag0.75%Cu3%Bi 224 10

2.4.3.1 Aerospace and Defense44,45

The screening experiments for the three alloys selected for the aerospace and defense

sector, as listed in Table 4, were performed between 2011 and 2013 by Celestica in

partnership with Honeywell. In these screening experiments, three alloys were compared

against baselines of both SAC305 and SnPb. Both BGA, using SAC305 solder balls,

except with SnPb solder in which case SnPb solder balls were used, and leaded

components typical of those used in aerospace and defense assemblies were investigated

using a medium complexity board designed by Honeywell, as shown in Figure 18.

Figure 18: Honeywell test vehicle, medium complexity board

27

2.4.3.1.1 Screening Experiment Structure

A Latin squares approach was used in this screening experiment.46

This approach allowed

for the most information to be obtained without testing every possible combination,

which would be cost prohibitive. Three factors were identified: board material, board

finish: Organic Solderability Preservative (OSP), Electroless Nickel Immersion Gold

(ENIG) and Electroless Nickel Electroless Palladium Immersion Gold (ENEPIG) and

paste (Violet, Orchid and Paul), in nine combinations rather than the full 27

combinations. Since only two Tg levels were of interest, instead of three, as with the

other factors, the higher Tg was repeated, as summarized in Table 5. Boards with Tg of

150C and 170C were used. The overall build matrix is summarized in Table 6. No

clean solder paste of the experimental alloys was prepared by a major solder paste

supplier for the purposes of this experimental work (Table 6). Boards were built at

Celestica. Two different reflow temperatures were used in the manufacturing process.

One with a peak temperature was 240C, was used with the SAC305 alloy and one with a

peak temperature of 222C for all other alloys. Both had a time above liquidus of

approximately 70-90 seconds.

Table 5: Latin square of paste and finish, and board materials

Alloy Finish

OSP ENIG ENEPIG

Paul 150 170 170

Violet 170 150 170

Orchid 170 170 150

28

Table 6: Build matrix showing number of assemblies

PWB Material Alloy Surface Finish

ENEPIG ENIG OSP

Normal Tg

(150 °C)

SnPb 2 3

SAC305 2 2

Paul 4

Violet 4

Orchid 4

High Tg

(170 °C)

SnPb 2 3

SAC305 2 1 3

Paul 4 4

Violet 4 4

Orchid 4 4

2.4.3.1.2 Microstructural Assessment

A detailed discussion on the microstructural evaluation can be found in Snugovsky et

al.44

and in Juarez et al.45

. All of the assemblies were inspected optically and with x-ray

radiography, to ensure that no major anomalies or concerns existed prior to further

testing. Figure 19 illustrates wetting of the solder paste on quad flat pack (QFP) leaded

components on OSP board finish. SnPb exhibited the best wetting properties, while

SAC305 showed the worst. All three experimental alloys showed some degree of

improved wetting over SAC305. Figure 20 shows the cross section evaluation performed

on solder joints made with all solder pastes on OSP board material. All BGAs showed

good mixing characteristics between the solder paste and the SAC305 solder ball.

Figure 19: Example of wetting on OSP, QFP240

29

Figure 20: BGA and QFP solder joints on OSP

Table 7 summarizes the composition of the resulting interconnect BGA solder ball after

assembly. This final BGA solder ball composition is a mixture of the paste alloy and the

original component SAC305 solder ball. The compositional measurements were made

using the semi-quantitative method of SEM-EDX and are provided here for comparative

purposes only.

Table 7: Resultant BGA interconnect composition

Paste

Alloy

Surface

Finish

Ball Composition (wt%)

Sn Ag Cu Bi Pb

SnPb OSP 70.2 0 0.4 0 29.4

SAC305 OSP 96.2 2.8 1.0 0 0

Paul OSP 95.0 2.6 0.8 1.7 0

Violet OSP 95.7 2.1 0.7 1.5 0

Orchid OSP 94.7 2.3 1.0 2.0 0

SAC305 ENIG 96.7 2.8 0.5 0 0

Paul ENIG 95.9 3.0 0.3 0.8 0

Violet ENIG 95.5 2.2 0.6 1.7 0

Orchid ENIG 95.0 2.5 0.4 2.1 0

Paul ENEPIG 95.0 2.5 0.4 2.1 0

Violet ENEPIG 95.1 2.8 0.6 1.5 0

Orchid ENEPIG 95.5 2.2 0.4 1.9 0

2.4.3.1.3 Accelerated Thermal Cycling (ATC)

17 assemblies of various paste, board finish, and board material combinations, were

exposed to upwards of 3000 cycles of harsh temperature thermal cycling in accordance

30

with IPC-9701A: Performance Test Methods and Qualification Requirements for Surface

Mount Solder Attachments47

using Test Condition 1 (-55C to 125C) for 3000 cycles.

Table 8 and Table 10 summarize the number of failures over the course of the entire test.

Mean Time to Failure (MTTF) is defined by Equation 1-5 and the failure rate () is

defined by Equation 1-6.48

Ni

i

iTN

MTTF1

1 2-5

Ni

i

iT

N

MTTF

1

1 2-6

Table 9 and Table 11 attempts to quantify the time to failure over the course of the 3,000

plus cycles in spite of the lack of a full statistical set of test data.

Table 8: ATC failures on 170C Tg board material

Paste type SnPb SAC305 Paul Violet Orchid

Finish ENEPI

G OSP

ENEPI

G OSP

ENEPI

G

ENI

G

ENEP

IG OSP ENIG OSP

352 BGA 2/2 2/2

240 QFP

31

Table 9: Reliability results on 170C Tg boards after ATC

Alloy Finish Component MTTF R(3039) =Failure/cycle

SnPb

ENEPIG 352 BGA - 1 0

240 QFP - 1 0

OSP 352 BGA 1611 0 0.00062

240 QFP - 1 0

SAC305


240 QFP - 1 0

OSP 352 BGA - 1 0

240 QFP - 1 0

Paul


240 QFP - 1 0

ENIG 352 BGA - 1 0

240 QFP - 1 0

Violet


240 QFP - 1 0

OSP 352 BGA 2271 0 0.00044

240 QFP - 1 0

Orchid

ENIG 352 BGA - 1 0

240 QFP - 1 0

OSP 352 BGA - 1 0

240 QFP - 1 0

Table 10: ATC failures on 150C Tg board material

Solder Alloy SnPb SAC305 Paul Violet Orchid

Finish ENEPIG OSP ENEPIG OSP OSP ENIG ENEPIG

352 BGA 2/2 2/2 2/2 2/2 1/2

240 QFP 1/2 2/2 2/2 2/2 1/2 1/2

32

Table 11: Reliability results on 150C Tg boards after ATC

Alloy Finish Component MTTF R(3039) =Failure/cycle

SnPb

ENEPIG 352 BGA 2310 0 0.00043

240 QFP 4466 0.5 0.00022

OSP 352 BGA 1336 0 0.00075

240 QFP 1313 0 0.00076

SAC305


240 QFP - 1 0

OSP 352 BGA 551 0 0.00182

240 QFP 802 0 0.00125

Paul OSP 352 BGA 1055 0 0.00095

240 QFP 1169 0 0.00086

Violet ENIG 352 BGA 4988 0.5 0.00020

240 QFP 5748 0.5 0.00017

Orchid ENEPIG 352 BGA - 1 0

240 QFP 4536 0.5 0.00022

In the above ATC testing, there are only two replicates of each condition. Even if both

fail, there is insufficient data to plot failure rates using Weibull, or any other type of

statistical distribution. Accordingly, MTTF and failure rate () were used to quantify the

data from this screening experiment and to describe the length of time for which each

condition survived. All combinations using the 170C Tg boards survived beyond the

1000 cycles required by the Aerospace and Defense sector. SnPb and Violet BGAs both

experienced failures with a MTTF of 1611 and 2271 respectively on OSP boards. As

OSP is not a finish that is favored by this sector, the results on the ENIG and ENEPIG

finish boards, which survived in all cases, show promise.

Combinations on the 150C Tg boards experienced more failures on all board finishes.

Only the SAC305 combinations on OSP show MTTF below the required 1000 cycles:

551 for BGAs and 802 for QFPs. Again, the boards with OSP finish failed earlier than

boards with other finishes. Upon further investigation, these failures were found to be in

the board material, in the form of barrel cracks, rather than in the solder joint

interconnect.

33

2.4.3.1.4 Vibration

Seventeen (17) assemblies of various paste, board finish and board material combinations

were exposed to two force levels of vibration testing. The test procedure, a discussion of

the failure modes, as well as a detailed discussion of the reliability results, can be found

in Juarez et al.45

Vibration Level (G)

Alloy

52

SAC3

05

Orch

idPa

ul

SnPb

Violet

Orchid

Paul

Violet

50

40

30

20

10

0

% F

ailu

re

Chart of % Failure

Figure 21: Failures in vibration testing

Figure 21 provides a summary of the number of failures that occurred during vibration

testing, at two different levels of acceleration, 2G and 5G respectively. The table

illustrates that all three experimental alloys outperform SAC305, with Violet even

outperforming SnPb. Table 12 and Table 13 summarize the failures that occurred based

on varying conditions and test combinations. A more comprehensive discussion of the

survival times can be found in the original work referenced above.

Table 12: Vibration failures after 2G testing

Alloy Paul Violet Orchid

Finish OSP ENIG ENEPIG OSP ENIG ENEPIG OSP ENIG ENEPIG

Board Material 150 170 170 170 150 170 170 170 150

352 BGA

240 QFP 2/2 2/2 2/2 2/2 1/2

34

Table 13: Vibration failures after 5G testing

Alloy Paul Violet Orchid

Finish OSP ENIG ENEPIG OSP ENIG ENEPI

G OSP ENIG ENEPIG

Board Material 150 170 170 170 150 170 170 170 150

352 BGA 2/2 2/2 1/2 2/2 2/2 2/2 2/2 1/2

240 QFP 2/2 2/2 2/2 2/2 2/2 2/2 2/2 2/2

Alloy SAC305 SnPb

Finish OSP OSP ENEPIG ENEPIG OSP OSP ENEPIG ENEPIG

Board Material 150 170 150 170 150 170 150 170

352 BGA 2/2 2/2 2/2 2/2 2/2 2/2

240 QFP 2/2 2/2 2/2 2/2 2/2 2/2 2/2

2.4.3.1.5 Summary of Findings

The three alloys compared against SAC305 and SnPb in the Aerospace and Defense

Sector screening experiments showed acceptable manufacturability characteristics with

improved wetting and voiding over SAC305. Two of the alloys, Paul and Orchid, did not

form proper intermetallics with ENIG or ENEPIG board finishes while Violet did. This

is finding is attributed to the lack of Cu present in these alloys.

All three experimental alloys met the Aerospace qualification of 1000 cycles of harsh

thermal cycling. Only SAC305 on 150C Tg boards with OSP finish failed to meet the

criteria, however, the failures were determined to be via failures within the board. This

good performance is attributed to the fact that Bi particles evenly precipitate into the Sn

matrix during thermocycling, reducing grain coarsening and microstructural degradation.

In vibration testing, all experimental alloys showed an improved performance over

SAC305, both Violet and Paul performed equally or better than SnPb.

3 Objective of Thesis: Low Melt Solders for Consumer Sector

The screening experiments for the three alloys selected for the Consumer Sector, as listed

in Table 4 and Table 14, are the focus of this thesis. While SAC305 has been widely

35

adopted by the industry during the initial transition to Pb-free, desired improvements in

mechanical and thermal performance, as well as cost reductions, continue to drive

research into new alloys.

Figure 22 illustrates the process for developing requirements for a new replacement alloy

for the consumer sector. The figure outlines the main shortcomings of the current market

available Pb-free solution based on SAC alloys, namely the high cost and poor

drop/shock performance. These shortcomings, in all probability could be overcome by a

new alloy, which is both less stiff and has a lower process temperature. The lower

process temperature would allow for “normal” Tg board material – that which has been

used with SnPb solders for many years – to again be used, eliminating collateral failure

modes such as pad cratering and an increased susceptibility to moisture delamination.

Figure 22: New approach proposing to use solder with no or low-Ag content6

Any replacement alloy for the consumer sector should perform better than SAC305 with

respect to drop/shock performance and thermal cycling. Both tests are representative of

service conditions that would likely be experienced by any consumer electronic product.

Improved Drop/Shock

Performance Less Expensive

Less Stiff Lower Process Temperature

Normal Tg Board Material

No Pad Cratering

No Moisture

Delamination/Cracks

New Pb-Free Solder

36

The following accelerated test conditions are accepted by industry as representative of

life conditions for these products:

ATC in accordance with IPC-9701A Performance Test Methods and Qualification

Requirements for Surface Mount Solder Attachments47

Drop/Shock Testing in accordance with the Joint Electron Device Engineering

Council (JEDEC) Standard JESD22-B110A Subassembly Mechanical Shock Test49

Table 14: Consumer alloys under test

Alloy Composition Assembly Temperature

SAC305 Sn 3%Ag 0.5%Cu 240C

Senju M42 Sn 2%Ag 0.75%Cu 3%Bi 224C

Sunrise Sn 1%Ag 0.7%Cu 7%Bi 222C

Sunflower Sn 0.7%Cu 7%Bi 226C

This thesis presents the results of both ATC and Drop Shock testing of the alloys selected

for consumer electronics applications (Table 4 and Table 14). These alloys were also

tested for a lower probability of Sn whisker formation according to JESD22A121.

37

4 References

1 K. Sweatman, J. Read, T. Nishimura, and K. Nogita “The Effect of Microalloy

Additions on the Morphology and Growth of Interfacial Intermetallic in Low-Ag

and No-Ag Pb-Free Solders”, presented at SMTAI, Chicago, Il, 2011.

2 K-N. Tu “Copper Tin Reactions in Bulk Samples” in Solder Joint Technology :

Materials, Properties and Reliability New York, Springer, 2007, ch. 1, pp. 1-32

3 IPC-A-610E “Acceptability of Electronic Assemblies” April 2010.

4 Directive 2002/95/EC of the European Parliament and of the Council of 27 January

2003 on the restriction of the use of certain hazardous substances in electrical and

electronic equipment

5 Directive 2011/65/EUof the European Parliament and of the Council of 8 June 2011 on

the restriction of the use of certain hazardous substances in electrical and


6 P. Snugovsky, S. Bagheri, M. Romansky, D. Perovic, L. Snugovsky, J. Rutter “New

Generation of Pb-Free Solder Alloys: Possible Solution to Solve Current Issues

with Main Stream Pb-Free Soldering,” SMTA Journal Volume 25 Issue 3, 2012.

7 M. Kelly, D. Colnago, V. Sirtori, C. Grosskopf, K. Lyjak, C. Ravenelle, E. Kobeda, J.

Bath, S.K. Tan, and L.H. Teo “Component Temperature Study on Tin-Lead and

Lead-Free Assemblies”, Journal of SMT, Volume 15 Issue 4, 2002.

8 IPC/JEDEC J-STD-020D.1 “Moisture/Reflow Sensitivity Classification for

Nonhermetic Solid State Surface Mount Devices” March 2008.

9 B.Gray “Correlation of Printed Circuit Board Properties to Pad-Crate Defects Under

Monotonic Spherical Bend”, M.A.Sc. thesis, Dept. Mech. Eng., Ryerson

University, Toronto, Ontario. 2012.

10 P. Snugovsky “Low Melt Pb-Free Solder to Solve Pb-Free Transition Challenges”

11 T. Sawamura and T. Igarashi “Difference between Various Sn/Ag/Cu Solder

Compositions”, Almit Ltd. June 2005.

12 K.S. Kim, S.H. Huh, K. Suganuma “Effects of intermetallic compound on properties of

Sn-Ag-Cu lead free solder joints” Journal of Alloys and Compounds 1 (2002)

13 L. Turbini. “Process and material issues related to Lead-free soldering”, J Mater Sci:

Mater Electron (2007) 18:147-154

14 U. Kattner, “Phase Diagrams for Lead-Free Solder Alloys” JOM December 2002 pp.

45-51

15 http://www.metallurgy.nist.gov/phase/solder/cusn-w.jpg

38

16 C. Handwerker. “Fundamental Properties of Pb-Free Solder Alloys”, in Pb-Free

Soldering, Springer 2007, ch.2, pp. 21-74

17 K-N. Tu “Introduction” in Solder Joint Technology : Materials, Properties and

Reliability New York, Springer, 2007, sec. 1.2.1, pp. 6

18 J. Dutkiewicz, L.A. Zabdyr, Z. Moser and J. Salawa. ASM Handbook Volume 3 Alloy

Phase Diagrams Materials Park, OH: ASM International 1992 pp. 2129

19 K.-W. Moon, W.J. Boettinger, U.R. Kattner, C.A. Handwerker, and D-J. Lee. “The

Effect of Pb Contamination on the Solidification Behavior of Sn-Bi Solders”,

Journal of Electronic Materials, Vol. 30, No. 1, 2001

20 M. Ribas S. Chegudi, A. Kumar, S. Mukherjee, S. Sarkar, R. Pandher, R. Raut, and B.

Singh. “Low Temperature Alloy Development For Electronics Assembly – Part

II”

21 P. Vianco, J. Rejent and R. Grant “Development of Sn-Based, Low Melting

Temperature Pb-Free Solder Alloys” Materials Transactions, Vol 45. No. 3

(2004) pp. 765-775

22 C. Handwerker. “Fundamental Properties of Pb-Free Solder Alloys”, in Pb-Free

Soldering, Springer 2007, ch.2, pp. 21-74

23 http://www.metallurgy.nist.gov/phase/solder/agcusn-l.jpg

24 http://www.empf.org/empfasis/2008/mar08/pillow.html

25 Acceptability of Electronic Assemblies IPC-A-610 Revision D February 2005 pp. 8-17

26 D.A. Shnawah, M.F.M. Sabri, I.A. Badruddin “A review on thermal cycling and drop

impact reliability of SAC solder joint in portable electronics products,”

Department of Mechanical Engineering, University of Malaya, Kuala Lumpur,

Microelectronics Reliability 52 (2012) pp. 90-99

27 P.T. Vianco, J.A. Rejent “Properties of Ternary Sn-Ag-Bi Solder Alloys: Part 1 –

Thermal Properties and Microstructural Analysis”, Journal of Electronic

Materials, Vol.28, No.10, pp. 1127-1137, 1999

28 P.T. Vianco, J.A. Rejent “Properties of Ternary Sn-Ag-Bi Solder Alloys: Part 2 –

Wettability and Mechanical Properties Analysis”, Journal of Electronic Materials,

Vol.28, No.10, pp. 1138-1143, 1999

29 http://www.metallurgy.nist.gov/phase/solder/bisn-w.jpg

30 http://www.metallurgy.nist.gov/phase/solder/agbisn.html

31 D.A. Porter and K.E.Easterling, “Thermodynamics and Phase Diagrams” in Phase

Transformations in Metals and Alloys, 2nd

ed. UK: Chapman & Hall, 1992, ch. 1,

pp 48-55.

32 http://www.metallurgy.nist.gov/phase/solder/bicusn.html

39

33 http://www1.asminternational.org/asmenterprise/APD/SearchAPD.aspx

34 J.S. Hwang “A Strong Lead-free Candidate: the Sn/Ag/Cu/Bi System”, SMT

Magazine, August 1, 2000.

35 http://www.senju.com/images/pdf/Senju%20ECO%20Solder.pdf

36 S.W. Chen, C-C Lin, C. Chen “Determination of the Melting and Solidification

Characteristics of Solders Using Differential Scanning Calorimetry”,

Metallurgical and Materials Transactions A, Volume 29A, July 1998 pp. 1965-

1972

37 L. Rycerz. “Practical remarks concerning phase diagrams determination on the basis of

differential scanning calorimetry measurements” J Therm Anal Calorim, 113,

2013 pp. 231-238

38 M. H. Kaye, K.M. Jaansalu and W.T. Thompson “Condensed Phases in Inorganic

Materials: Metallic Systems” in Measurement of the Thermodynamic Properties

of Multiple Phases, R.D. Weir and Th.W. deLoos, Ed. San Diego, Elsevier Inc.,

2005, pp. 227-305

39 J.S. Hwang “A Strong Lead-free Candidate: the Sn/Ag/Cu/Bi System”, SMT

Magazine, August 1, 2000.

40 http://www.senju.com/images/pdf/Senju%20ECO%20Solder.pdf

41 J.W. Evans “Introduction to Solder Alloys and Their Properties” in A Guide to Lead-

free Solders: Physical Metallurgy and Reliability, Silver Spring, MD: Springer,

2005, pp. 1-27

42 J.W. Evans “Microstructural Instability in Solders” in A Guide to Lead-free Solders:

Physical Metallurgy and Reliability, Silver Spring, MD: Springer, 2005, pp. 79-

95

43 E. Kosiba S. Bagheri, Z. Bagheri, P. Snugovsky, and D. Perovic “Assembly Feasibility

and Property Evaluation of Low Ag, Bi-Containing Solder Alloys” in ICSR

SMTA Conference, Toronto, ON, 2012

44 P. Snugovsky, E. Kosiba, J. Kennedy, Z. Bagheri, M. Romansky, M. Robinson, J.M.

Juarez, Jr., J.Heebink “Manufacturability and Reliability Screening of Lower

Melting Point Pb-free Alloys Containing Bi,” in IPC APEX EXPO Conference,

San Deigo, CA, 2013.

45 J. Juarez Jr., M. Robinson, J. Heebink, P. Snugovsky, E. Kosiba, J. Kennedy, Z.

Bagheri, S. Suthakaran, M. Romansky “Reliability Screening of Lower Melting

Point Pb-Free Alloys Containing Bi,” in IPC APEX EXPO Conference, Las

Vegas, NV, 2014.

46 D.C. Montgomery. “Design and Analysis of Experiments”, 8

th Edition. Jonh Wiley &

Sons, 2012.

40

47 IPC-9701A Performance Test Methods and Qualification Requirements for Surface

Mount Solder Attachments, February 2006

48 J. Bentley “Introduction to Reliability and Quality Engineering”, 2

nd Ed, Essex,

England, Pearson Education Limited, 1999, ch. 2, pp 28-43

49 Subassembly Mechanical Shock, JESD22-B110A, November 2004.

41

Chapter 2 Solder Joints after Reflow (As Manufactured)

1 Introduction

The formation of a solder joint is influenced by the conditions at which the chemical

bond is formed: primarily, the maximum temperature of the reflow, the time above

solidus and the cooling rate. Therefore, the melting temperature and pasty range are

primary characteristics which differentiate solder alloys. The solder joint microstructure

further evolves when exposed to elevated temperatures for a prolonged period of time or

if exposed to other forms of stress, due to solid-state diffusion and other physical effects

described below. The microstructural properties described in this chapter, which result

from the solidification of liquid solders, will have an influence on the mechanical and

thermo-mechanical properties of the solder joint and therefore impact the reliability in

field conditions.

2 Bulk Solder Microstructure – Solidification Process

Slight changes to the alloy composition may result in different solidification regimes

within the solder joint and therefore significantly change properties. It is therefore of

interest to examine the factors which govern this process.

2.1 Cu Dissolution in Molten Solder

As the molten solder comes into contact with the solid Cu pad layer on the PWB board,

Cu will dissolve into the liquid solder. For this reason, the overall concentration of Cu

within the molten solder will increase at a rate that is dependant on the time spent above

the solidus temperature. SAC305, as an example, has an initial Cu concentration of

0.5wt% however, by the time cooling is initiated, the Cu concentration may be as high as

1wt%. Figure 23 shows a typical SAC305 reflow profile in which a particular solder joint

experienced 71 seconds above 217C, the solidus temperature for SAC305. This Cu

dissolution during the liquid phase of solder joint formation will continue, given enough

42

time, until the liquids solubility limit is reached. This reaction has been characterized by

the Nernst-Brunner equation (Equation 2-1):

V

KAt

XX

XXS

s

0

ln 2-1

where XS

represents the solubility limit, X0 represents the initial concentration, K is the

temperature-dependent Nernst-Brunner dissolution rate constant, A is the contact area at

the interface, V is the total volume of solder, t is time.1

Figure 23: Typical reflow profile for SAC305

If dissolution kinetic conditions are satisfied (for example if a very high temperature is

maintained over a sufficient length of time), the Cu present in the board material can be

dissolved into the solder up to a very high concentration. Potentially the entire copper pad

region on the PWB can be consumed by the intermetallic and bulk solder. Figure 24

shows the Cu pad being progressively consumed by the IMC layer after successive

rework/replacement cycles. In this example, the solder joint is repeatedly heated above

the solder alloy liquidus temperature.

43

Figure 24: Copper dissolution of BGA solder joint after a) 1 replacement b) 3

replacements and c) 5 replacements2

It has been shown that an increase in copper concentration within a Sn-based solder alloy

from 0.5%Cu to 0.7%Cu decreases the amount of copper dissolution from the PWB

material, which is attributed to the reduction in the concentration gradient.3 This is the

primary reason for the presence of Cu in the SAC system of alloys and in the alloys

proposed in this study. The Cu dissolution described above typically results in a

concentration within the hypereutectic region of the Cu-Sn phase diagram (Figure 6). As

the solder cools Cu6Sn5 particles will begin to nucleate within the liquid.

As the Cu dissolution into the liquid solder occurs, a competing process is taking place in

which the Cu and liquid solder react to form a chemical bond at the interface, which takes

the form of an intermetallic compound – the interfacial IMC. This reaction is discussed in

3.

2.2 Bulk Solder Solidification

Within the bulk solder, the equilibrium solidification of SAC305 follows the path

outlined in red, in Figure 25, where non-faceted Sn grains begin to form a dendritic

structure. As the liquid continues to be consumed by the new Sn structure, the

remaining liquid composition moves along the red line from A towards B. At B, a

eutectic of Sn + Ag3Sn forms in the interdendritic spaces of the already solidified Sn

dendrite. This eutectic solidification continues, and the liquid compositions following the

eutectic valley from B towards E. At this point, any remaining liquid is now at the ternary

44

equilibrium composition (point E) and will solidify as a ternary eutectic of Sn + Ag3Sn

+ Cu6Sn5.4

Figure 25: SnAgCu phase diagram with SAC305 equilibrium solidification path5

The solidification path described above does not account for Cu dissolving into the bulk

solder from the board side Cu pad and, in the case of a leaded component, from the lead

itself. The Cu concentration in the bulk liquid is therefore increasing; the blue line in

Figure 25 indicates a concentration of approximately 1wt% Cu. In this case, the Cu6Sn5

begins to solidify before the Sn. The Cu6Sn5 continues to form until the concentration

reaches the eutectic line and a binary eutectic of Sn + Cu6Sn5 forms. Additionally,

Cu6Sn5 IMC, which has already formed at the Cu-liquid solder interface, and will be

described in section 3, may break off and move through the bulk solder.

As the Sn dendritic arms grow, the composition of the liquid at the Sn/liquid interface

changes and becomes Cu rich. In the Cu rich regions, a binary of Sn + Cu6Sn5 will

begin to nucleate and grow. The remaining liquid at the Sn + Cu6Sn5 eutectic/liquid

interface will be Ag rich. At this point, the remaining liquid will either solidify as a

45

eutectic of Sn + Ag3Sn to form the “double binary eutectic”6, or as a ternary eutectic of

Sn + Cu6Sn5 + Ag3Sn.

2.3 Undercooling During Solidification

The solidification path described above assumes equilibrium cooling conditions, in which

sufficient time is allowed for all processes to occur. In reality, cooling rates are usually

faster than equilibrium conditions resulting in some effects of undercooling. A number

of different structural morphologies can be expected with different cooling rates. For

example, a quick quench after melting of a SAC alloy resulted in a fine Sn dendrite

structure with very fine, non-faceted ternary eutectic structure in the interdendritic

spaces.7

Figure 23 shows that in a typical SMT reflow process, the solder alloy cools relatively

quickly after reaching maximum temperature. This cooling profile will induce

microstructures more typical of undercooling conditions than of equilibrium cooling as

described above. In the case of a SAC alloy, in which Sn, Cu6Sn5 and Ag3Sn all form,

cooling below the eutectic temperature of 217C would represent an undercooling

condition for all three of the solid phases. As one of these phases begins to nucleate and

grow, the composition of the remaining liquid changes. A move away from the eutectic

point would represent a condition of even further undercooling of one of the phases, as,

in all cases the liquidus surface increases away from the eutectic point. As the second

phase nucleates, the remaining liquid again shifts in concentration until the nucleation of

the third and final phase occurs. This is referred to as constitutional undercooling,

describing the changing conditions of the liquid at the solidification front.

As described above, it is expected that a SAC305 solder, cooled quickly below the

ternary eutectic point of 217C, would initially form a Sn phase, followed by Sn +

Ag3Sn5 eutectic and finally the ternary Sn + Ag3Sn5 + Cu6Sn5 eutectic. As the

concentration of Cu in the molten solder is increased, Cu6Sn5 is favored to solidify first.

This implies that the primary phase to solidify will be the one with the highest liquidus

temperature, however there are a number of other factors to consider including:

46

nucleation from undercooled liquid

the growth kinetics of facetted vs. non-facetted phases

the availability of Cu and Ag within the volume to form corresponding phases.6

Figure 26 shows a 3D phase diagram in the Sn rich portion of a SAC alloy. The liquidus

surfaces are extended below the eutectic point to help visualize the solidification of the

three phases in undercooling situations. The red line shows the equilibrium solidification

of a particular Ag rich alloy, while the white line shows the solidification path of the

same alloy in non-equilibrium, undercooling conditions. In this scenario, Ag3Sn

continues to form until the composition reaches the eutectic of Ag3Sn + Cu6Sn5.

The undercooling of SAC305 with further dissolved Cu (Cu concentration of ~1wt%)

would result in a solidification path along the Cu6Sn5 surface below the eutectic point.

As the temperature is significantly lower than the liquidus temperature for this phase,

Cu6Sn5 will nucleate and grow. The composition of the liquid will then change and move

to another area on the phase diagram. The next phase to solidify will be the one with the

highest liquidus temperature, and therefore the greatest degree of undercooling. This will

continue until all liquid has solidified.

Figure 26: A 3D phase diagram of Sn rich portion of Sn-Ag-Cu ternary system8

47

The growth kinetics of the three phases also plays a role in the solidification process.

While this process is beyond the scope of this work, it should be considered that both

Ag3Sn and Cu6Sn5 have facetted structures and are therefore restricted to grow in specific

crystallographic directions, while the non-facetted Sn can grow through the fluid in a

more unrestricted fashion.6

Finally, the volume fraction of elements in the liquid needs to be considered. As one

phase solidifies, its growth is further restricted by the availability of atoms of the correct

type. The final volume fraction of the Ag3Sn and Cu6Sn5 phase is therefore limited by

the availability of Ag and Cu atoms respectively within the liquid. Ag3Sn requires a

localized concentration of 73.2wt% Ag within a total volume with a concentration of only

3wt% in SAC305 and even lower in the other alloys considered in this study. This would

require significant diffusion of Ag through the liquid, which may not occur in

undercooling conditions. Similarly Cu6Sn5 requires a local concentration of 39.1wt% Cu

within a liquid with only approximately 1wt% Cu. Sn, by contrast, will likely have a

ready supply of Sn at the solidification fronts of the other two phases since it makes up

the vast majority of the volume fraction.

2.4 Formation of Facetted IMCs in Bulk Solder

The presence of large, faceted IMCs in the bulk solder, Ag3Sn and Cu6Sn5, may present a

reliability concern as they can act as stress risers and paths along which cracks

propagate.9 Large, primary Cu6Sn5 particles form when the total concentration of Cu in

the liquid is high and Cu6Sn5 is the first phase to solidify (Figure 25). Some undercooling

is also required, however the undercooling required to nucleate Cu6Sn5 is less than that

required to nucleate Sn.10

This phase has a facetted structure of hexagonal morphology,

often hollow on the inside. The structure then grows in a rod, forming a structure similar

to a pencil – a hexagonal tube with a facetted center (Figure 27). As a primary phase

solidifying from liquid, this facetted rod structure will be surrounded by a liquid whose

composition moving along the blue line shown in Figure 25 until a eutectic of Sn +

Cu6Sn5 forms together. Since the base phase (Cu6Sn5) is a more complex crystal

48

structure, it dictates the final shape of the eutectic colony. In this case, the Cu6Sn5 portion

of the eutectic may form as branches of the original rod like structure.11

Ag3Sn IMCs form within the bulk solder. Similar to the Cu6Sn5 structures, Ag3Sn may

form as primary phases, or as part of a binary or ternary eutectic solidification regime

depending on the alloy composition and cooling rate. At compositions below

approximately 3.7wt% Ag, Ag3Sn particles are not expected to form as a primary phase.

It is more likely that the Ag3Sn will form in the interdendritic spaces of the Sn and/or

Cu6Sn5 phases and as part of a eutectic with one of both of the other two phases. In a

typical cross-section view, the Ag3Sn IMC will appear as small particles. In fact, these

particles have a fine, fibrous structure (Figure 27), which can be seen after selectively

etching away the Sn phase. The network of fibers typically appears to surround the Sn

dendrite arms, indicating that they form in the interdendritic spaces during eutectic

solidification. 12

Figure 27: SAC solder joint viewed a) in cross section and b) after selective

electrochemical etching to remove Sn phase13

2.5 Bi in Solution and as a Precipitate

As discussed in Section 2.4.1.2, the three experimental alloys shown in Table 15 were

selected from outside the ternary eutectic triangles of the ternary Sn-Ag-Bi and Sn-Bi-Cu

alloys so as not to form any ternary eutectic phase with Bi. This is both to avoid the

49

formation of a low melting temperature (~138C) phase, as well as to avoid the formation

of primary Bi particles, which form directly from the liquid state. Depending on the

amount of Bi in the overall composition, favourable Bi particles may be produced in the

final solder joint microstructure; however these particles are the result of solid solution

precipitation from within the Sn. Bi precipitation from solid solution has been

confirmed by the DSC curve for each studied alloy. In all cases, only one main peak

appears, indicating that all liquid was consumed prior to the ternary eutectic solidification

which could have resulted in a primary Bi phase.

Bi, as an alloying element in a Sn based solder, differs from Ag and Cu in two significant

ways: Bi has high solubility in Sn (up to 21wt%), and Bi does not form any intermetallics

with Sn. It should also be noted that Sn is not soluble in Bi; any Bi particles that form

will therefore be 100% Bi.

3 Sn-Cu Reaction and Formation of Interfacial IMC

The chemical bond which forms between Sn and Cu provides a strong interaction

between the solder and the substrate. As described earlier, this bond provides a chemical,

mechanical and electrical bond between the two materials. Both Cu6Sn5 () and Cu3Sn

() phases can potentially form during solidification at the interface. It has been shown

that the formation of the IMC layer increases with increased temperature.14

This

generally occurs up to a thickness of approximately 2.5µm at which point the IMC

appears to form as more needle like faceted rods, initially at the interface, and then

through the bulk solder (Figure 28). The phase was not shown to form at initial

solidification, but rather occurred over time as an effect of aging. This will be discussed

further in Chapter Chapter 3.

Many studies looking at the Sn-Cu diffusion couples considered two solid substrates in

contact while exposed to heat. When considering a solid Cu with a molten Sn coupling

to form a molten-Sn/solid-/solid-Cu interface, two rate dependent steps need to be

considered:14

50

Diffusion of Sn through to solid Cu interface

Reaction of Cu and Sn at this interface

Figure 28: Solder joint between Sn-based solder and Cu a) rod shaped Cu6Sn5 ()

b) scallop shaped Cu6Sn5 () IMC layer

SAC305 has solidus and liquidus temperatures of 217C and 220C respectively. A

typical reflow profile (Figure 23) will have a time above solidus of anywhere from 60 to

90 seconds and a maximum temperature of approximately 240C. The initial formation

of the interfacial IMC layer between the Cu and phase is governed by time, temperature

and contact area. While the contact area is fixed for a particular build, both the time and

temperature can be varied by altering the solder reflow profile or solder alloy

composition.

Although covered in Chapter Chapter 3, it is worth noting that the interfacial IMC will

continue to grow at rates determined by time and temperature. At lower aging

temperatures (ambient temperature to about 70C), the layer grows preferentially

indicating that the Cu diffuses at these temperatures, while at higher storage temperatures

(above 135C) both the and layer will form and grow, indicating that Sn diffusion

begins to dominate at these increased temperatures.15

Another possibility is that the Sn

supply between the phase and the copper interface is inadequate for the formation of

the phase.16

The morphology of the /solder interface in a typical solder joint tends to

be scalloped in shape (Figure 28b) but may flatten to a more planar interface as it grows

51

upon aging. The phase, which is sandwiched between the and the Cu (resulting in a

//Cu interfaces), tends to form a more planar morphology.16,19

It has also been observed that microvoids may form between the copper and interface17

.

These Kirkendall voids18

form as the Cu diffuses into the layer leaving behind voids

(Figure 29a). These present a long-term reliability concern as they weaken the interface.

Another consideration is the morphology of the interface that is formed.

In the SnPb solder system joined to a copper substrate, the Sn is consumed from the area

close to the solid copper forming the Cu6Sn5 intermetallic compound. This can leave a

localized region of Pb in direct contact with the intermetallic (Figure 29b). This effect

may continue under aging conditions as the Cu6Sn5 continues to grow, and the Sn in the

immediate vicinity is depleted.19

This presents a potential reliability concern; the Cu6Sn5

being brittle and the segregated Pb providing a convenient path for crack propagation and

eventual fracture. This combination has been observed to fracture predominately along

the /solder interface.19

Figure 29: SnPb solder jointed to Cu substrate a) showing Kirkendall voids20

b)

showing Pb phase pooling at IMC/solder interface (cross section by Zohreh Bagheri)

In this manner, Bi has been found to behave in a similar way as Pb. SnBi eutectic solders

may therefore have the same reliability concerns related to the pooling of one phase, Bi,

along the Cu6Sn5 interfacial IMC. The lower level of Bi used in this study, between 3-

7wt%, is not expected to result in this condition.

52

The scallop shaped morphology, typically associated with the Cu6Sn5 IMC bond layer, is

a result of the solidification of the molten solder and dissolved Cu6Sn5 at the interface. A

fast, non-equilibrium cooling rate typical in a solder reflow profile will result in a

scalloped morphology. A slower rate will allow for the interfacial IMC layer to form in a

smoother morphology and a faster rate will result in a more complicated dendritic, or

“coral” shaped structure. It has also been shown that the structure of the base Cu is a

contributing factor to the morphology of the interfacial IMC layer.21

It was found that

Cu6Sn5 grows in a prism-type morphology from single crystal Cu with (001) or (111)

grain orientations (Figure 30). Other crystallographic orientations result in a scallop-type

structure similar to that found in a polycrystalline Cu substrate. Since the board side Cu

pad is polycrystalline and not single crystal, it will form a scallop shaped IMC layer.

However, in localized areas where a Cu grain may have a (001) or (111) orientation, a

prism structure may form amongst smooth scallops.

Figure 30: Typical morphologies of Cu6Sn5 grains formed on single crystal

a) (001) Cu and b) (011) Cu21

Growth of the IMC bonding layers can also occur by solid-state diffusion during aging

conditions. This will be considered in Chapter Chapter 3.

53

4 Cu-Ni-Sn Interface

Because of the tendency for Cu to oxidize, protective layers on the Cu surfaces of PWBs

are used. For this project an OSP finish was used which is dissolved by the flux in the

solder paste during the reflow, and does not need to be considered when evaluating the

resulting solder joint. Other board finishes utilize layers of Ni, Au, Pd or other inert

metals as protective barriers. These are usually present in very thin layers, often in

combination. For example, an ENIG finish is made up of a thin Au layer (0.05µm) over

an electroless Ni layer (3-6µm). Additionally, it is common to find an electrolytic Ni

barrier layer used between the Cu pad on a BGA component and the attached solder ball;

it is for this reason that the Cu-Ni-Sn interaction is important in solder joints formed in

this study. Ni also has a much slower diffusion rate than Cu in molten solder therefore it

serves as a diffusion barrier to Cu. With a SnPb solder, it was found that a scallop shaped

Ni3Sn4 intermetallic formed, however with SAC alloys this phase appeared to be

suppressed in favour of a (Cu,Ni)6Sn5 IMC bonding layer. It is the Cu from within the

solder which is present in the final IMC composition.22

This is verified by examining a

similar IMC formed using a SnPb solder, in which no Cu is present; it forms a Ni6Sn5

IMC with no Cu present. The (Cu,Ni)6Sn5 phase is made up of many needle-like,

hexagonal cylinders with pointy tips or, when seen as a layer has a “coral” type

morphology (Figure 31).23,24

Snugovsky et al. 25

showed that in the Sn rich corner of the Cu-Ni-Sn system, two quasi-

peritectic reactions exist:

L+(Ni,Cu)3Sn4 ↔ Ni23Cu33Sn44+Sn at 229.1C 4-1

L+Ni23Cu33Sn44 ↔ (Cu,Ni)6Sn5+Sn at 228.1C 4-2

Therefore the IMC bond layer formed at the component side of the BGA, which is

formed from the Ni barrier layer, and both Sn and Cu from the SAC305 alloy will likely

be one of the compounds listed in Equation 2-2 and 2-3. Additionally, the IMC which

forms towards the board side of the BGA component will be formed primarily from the

Cu layer and the Sn from within the SAC305 solder ball and paste. Ni, which has

54

diffused through the molten solder, may also be present. The Ni23Cu33Sn44 intermetallic

has a smooth, cellular morphology. 26

The formation of Kirkendall voids is also a concern

at the (Cu,Ni)6Sn5 – Ni(P) interface.

Figure 31: IMC formed between SAC solder and Ni(P)Au23

5 Experimental Set Up

5.1 Test Vehicle Assembly

Celestica’s RIA3 test vehicle (Figure 32) was used for the work described in this paper. It

was originally designed to simulate a typical, medium complexity assembly. It is an

8”x10” surface made up of 12 Cu layers for a total thickness of 0.093” with an OSP

finish. This board is one which is often used to test new, lead-free solders and other

material parameters. LQFP176, PBGA256, CBGA64 and MLF20 components were

populated, two of each on each board. The BGA components all employed SAC305 ball

alloy. The board and SMT components were attached with solder pastes of the alloys of

interest. The solder paste alloys were experimental formulations of the alloy and paste

flux suitable for the process temperatures. The SAC305 and Senju M42 pastes were

commercially formulated by Indium Corp. The reflow was performed in an air

environment in a 10 zone oven. One board from each combination listed in Table 15 was

held for evaluation of the “As Assembled” microstructure.

55

Figure 32: Test vehicle

Table 15: Build matrix for as assembled analysis


SAC305 Sn 3%Ag 0.5%Cu 240C




5.2 Test Matrix

Each of the four components listed in Table 16 was cross sectioned to evaluate the solder

mixing, shape of the solder joint, degree of voiding and any possible anomalies.

Table 16: Table of components evaluated by cross sectioning

Component Reference Designator

BGA-256 U204

BGA-256 U205

QFP-176 U1

QFP-176 U2

56

5.3 Test Method

Components were cut from the test vehicle shown in Figure 32 and mounted in epoxy for

cross-sectioning. Each cross-section was ground and polished through the following

sequence: 500 and 1200 grade SiC paper, polishing with 6 µm and 1 µm DiaPro diamond

suspensions (Struers), and an oxide polish (Struers OP-S). Optical microscopy was

performed using a Nikon Measurescope MM-11. Prior to SEM analysis, the samples

were carbon coated using an Emitech K950X. SEM microscopy was performed using a

Hitachi S-4500 and Hitachi S-3000N with the following EDX systems: Oxford and

ThermoScientific respectively.

6 Microstructural Evaluation

6.1 Comparison of Bulk Microstructure

EDX was used to evaluate both the overall composition of the bulk material in the solder

joint and the composition of Sn grains within the solder joint (Figure 33). Although EDX

is a semi-quantitative approach, it was deemed sufficient for purposes of comparison.

Scans of the bulk solder joints were taken of cross sectioned samples as seen in Figure 33

a) and b) as well as of the Sn grain (Figure 33c) for each alloy. The results are tabulated

in Table 17 and Table 18. Each value represents an average of 3-5 measurements.

Figure 33: Area for EDX compositional analysis of a) bulk BGA solder joint,

b) bulk QFP solder joint and c) Sn grain

57

6.1.1 Bulk Microstructure of QFP Solder Joints

Table 17: Composition of QFPs (wt%)

Paste Alloy

Composition of

Paste Alloy

(wt%)

Bulk Solder

(wt%) Sn Grain

(wt%)

Sn Ag Cu Bi Sn Bi

SAC305 Sn 3%Ag 0.5%Cu 96.8 2.2 1.0 0.0 100 0

Senju M42 Sn 2%Ag 0.75%Cu

3%Bi 94.5 1.8 1.2 2.5 98.0 2.0

Sunrise Sn 1%Ag 0.7%Cu 7%Bi 92.7 0.9 1.0 5.6 95.0 4.0

Sunflower Sn 0.7%Cu 7%Bi 93.8 0.0 1.0 5.2 94.5 5.5

The bulk composition of the QFP solder joints is close to that of the original paste

composition with the exception of increased Cu concentration and a very small increase

in the Sn content. The increase in Cu concentration is expected due to Cu dissolution

from the board side Cu pad and from the lead. The leads are plated with a thin layer of

Sn, which melts during the reflow process and combines with the bulk solder. It is also

expected that the concentration of Bi within the Sn grain would be lower than in the

overall composition as can be seen in two of the three Bi-containing alloys in Table 17,

Senju M42 and Sunrise. This is a result of the off-eutectic solidification in which primary

Sn dendrite arms form prior to the final eutectic solidification of the remaining liquid.

The primary Sn grain will continue to grow, driving the Bi content of the liquid at the

solid/liquid interface to increase. Finally, the remaining liquid solidifies as a eutectic

structure within the interdendritic spaces. The Bi concentration of the eutectic portion of

the Sn will therefore be greater than that of the primary Sn dendrite arms. Over time

and with the precipitation of Bi from solid solution, it is expected that this concentration

gradient will diminish. The Sunflower alloy has a composition which lies on the eutectic

line on the Sn-Bi-Cu phase diagram (Figure 13). In this case, a eutectic of Sn + Cu6Sn5

will solidify, instead of the nucleation and growth of Sn grains prior to solidification of

other phases. The Bi content of the Sn dendritic arms in Sunflower is similar to the

overall final composition of the solder joint, as seen in Table 17.

Figure 34a and b show the microstructures of SAC305 and Senju M42 solder joints

respectively. Both have very similar microstructures with fine Sn dendritic arms

58

surrounded by fine Ag3Sn and Cu6Sn5 IMCs solidified as either binary or ternary

eutectics with Sn in the interdendritic spaces. In the case of Senju M42, all Bi appears to

be in solid solution with Sn, with no visible Bi precipitates.

Figure 34c and d show cross sectional images of Sunrise and Sunflower respectively.

Both have larger Sn dendritic arms and Cu6Sn5 IMCs. Sunrise shows a small amount of

Ag3Sn particles, while Sunflower does not have any due to the lack of Ag in the paste

composition. In Sunrise, while Bi particles are visible and believed to have precipitated

from solid solution of the interdendritic Sn, which exists as either a binary eutectic with

Sn + Ag3Sn, Sn + Cu6Sn5, or a ternary eutectic of Sn + Ag3Sn + Cu6Sn5. As described

above, this interdentritic, eutectic Sn, rather than that found in the primary Sn dendrite

arms, has a higher Bi content resulting from the solidification of the off eutectic alloy

(Table 17). Sunrise shows an overall composition with 5.6wt% Bi, and only 4.0wt% in

the primary Sn dendrite arms. The bulk of the Bi, therefore is in the Sn within the

interdendritic, eutectic colonies (Figure 35a). Sunflower on the other hand shows the

presence of Bi in both interdendritic and within the Sn dendritic arms (Figure 35b).

59

Figure 34: Optical images QFP a) SAC305 b) Senju M42 c) Sunrise and d)

Sunflower

Figure 35: Bi present in a) eutectic colonies in Sunrise and b) throughout bulk

solder (eutectic colonies and Sn dendrite arms) of Sunflower

60

The presence of Bi, as well as the distinction between Cu6Sn5 and Ag3Sn, are visible

optically (Figure 36a) however are better distinguished in scanning electron images,

particularly in backscatter electron (BSE) images (Figure 36c) in which differences in

atomic numbers are more visible. In Figure 36b and c, the Bi is shown in light colour

corresponding to a higher atomic number (83) than the surrounding Sn (50). In the

secondary electron (SE) image (Figure 36b), Cu6Sn5 and Ag3Sn appear indistinguishable,

however in BSE (Figure 36c), Cu6Sn5 appears significantly darker while Ag3Sn becomes

difficult to distinguish from the Sn.

Figure 36: QFP solder joint as seen a) optically b) SE SEM and c) BSE SEM

The Bi precipitating from the interdendritic Sn appears in both Sunrise and Sunflower

as both large particles and as fine dispersions (Figure 37). It is expected that the Bi atoms

within the Sn matrix will diffuse together to form a small volume of Bi precipitate, and

then will rearrange themselves into the Bi crystal structure.27

Bi, unlike Cu6Sn5 and

Ag3Sn, has a non-faceted structure.

61

Figure 37: BSE images of a) Sunrise QFP solder joint and b) Sunflower QFP solder

joint

6.1.2 Bulk Microstructure of BGAs

The final composition of a BGA solder joint is made up of both the screened solder paste

as well as the component solder ball. The BGA-256, which was examined in this study,

is made up of SAC305 solder spheres with a diameter of 30 mil and solder paste of each

of the four alloys screened using a 5 mil stencil and a 30 mil aperture. This results in a

volume composition of approximately 87% SAC305 solder sphere and 13% solder paste.

Table 19 provides a summary of the expected final composition of the solder joint formed

from the original solder sphere and the screen solder paste. Table 18 provides a summary

of the composition as measured using EDX as shown in Figure 33a. As with the QFPs,

the composition is close to the expected composition with the exception of an increased

Cu concentration. This is the expected result due to Cu dissolution from the board side

Cu pad. The Cu content increase in SAC305 is greater than that of the other three alloys,

which can be attributed to the higher process temperature. Cu dissolution from the

component side of the solder joint would be mitigated by the Ni barrier layer described in

Section 4. Table 22 indicated that some Ni was present in the board side IMC. This

would indicate that Ni from the component side would diffuse through the solder during

liquid phase mixing and may be present in the final bulk solder composition. Ni was not

found to be present and may therefore be present in trace amounts only.

62

Table 18: Experimental composition of BGA solder joint

Paste

Alloy Composition of

Paste Alloy

(wt%)

Bulk Solder

(wt%) Sn Grain

(wt%)

Sn Ag Cu Bi Sn Bi

SAC305 Sn 3%Ag 0.5%Cu 95.5 2.7 1.8 0.0 100 0

Senju

M42

Sn 2%Ag 0.75%Cu

3%Bi 96.2 2.4 0.9 0.5 99.7 0.3

Sunrise Sn 1%Ag 0.7%Cu

7%Bi 95.6 2.6 0.8 1.0 99.3 0.7

Sunflower Sn 0.7%Cu 7%Bi 95.8 2.2 1.0 1.0 99.1 0.9

Table 19: Theoretical composition of BGA solder joint

Paste Alloy

Composition of

Paste Alloy

(wt%)

Theoretical Composition

of Bulk Solder

(wt%)

Sn Ag Cu Bi

SAC305 Sn 3%Ag 0.5%Cu 96.5 3.0 0.5 0.0


3%Bi 96.2 2.9 0.5 0.4

Sunrise Sn 1%Ag 0.7%Cu 7%Bi 95.8 2.7 0.5 0.9

Sunflower Sn 0.7%Cu 7%Bi 96.0 2.6 0.5 0.9

Table 19 gives the expected final composition of BGA solder joint based on the volume

of paste versus the SAC305 solder ball. This calculation does not take into account Cu

dissolution. Table 18 is the observed composition measured by EDX. Of note is the

increase in Cu content. Senju M42, Sunrise and Sunflower all have similar Cu

concentrations between 0.8-1.0wt%, which is also similar to that found in the QFPs. The

solder joint formed fully of SAC305 has almost tripled the Cu content at 1.8wt%. This is

likely due to the increase in process temperature; the boards built with SAC305 solder

paste reached a maximum temperature of 240C where as the other solder pastes reached

maximum process temperatures between 222 and 226C.

Good mixing of these two constituent materials is demonstrated in Figure 38 and Figure

39. These images show a portion of a BGA solder joint, formed between SAC305 solder

63

ball and Sunflower solder paste, close to the board side IMC. The EDX mapping shows

the presence of Ag at the lower end of the solder joint. As the screened solder paste,

Sunflower, contributed no Ag to the composition, this indicates good mixing from the

solder ball.

Figure 38: SEM image of Sunflower BGA solder joint

64

Figure 39: EDX mapping of intermetallic in Sunflower showing a) image

b) Sn c) Ag and d) Cu

In all solder joints shown in Figure 40, there are no visible precipitates of Bi found in any

of the BGAs solder joints formed. With an overall composition of up to 1wt% Bi, it is not

expected that Bi would precipitate from solid solution under equilibrium conditions.

BGA solder joints formed between SAC305 solder balls and each of the four alloys under

investigation all exhibited Cu6Sn5 flower shaped structures, which were a complex

dendritic structure of the Cu6Sn5 pencil shaped rods with branches of the same shape

(Figure 41). These are primary IMCs that form directly from liquid, indicating some

undercooling conditions. The equilibrium phase diagram of Sn-Ag-Cu (Figure 25)

indicates that Cu6Sn5 would be the first phase to solidify and continue for a short time

before the solidification of Sn + Cu6Sn5 eutectic. The presence of large and highly

65

branched Cu6Sn5 faceted structures indicates some degree of undercooling during the

solidification process. In this case, the Cu6Sn5 would continue to grow for an extended

period prior to the formation of other phases. These Cu6Sn5 flowers appeared in solder

joints formed with all four examined solder pastes.

Figure 40: Optical images BGA formed with SAC305 solder balls and paste with the

following alloys a) SAC305 b) Senju M42 c) Sunrise and d) Sunflower

66

Figure 41: Cu6Sn5 flower structures seen in BGA solder joints made up of SAC305

and a) SAC305 and b) Sunflower solder paste

Figure 42 shows the cross sectional image of a number of BGA solder joints under

polarized light. Polarized light is used to reveal the grain structure of the solder joint. As

is expected from this size of solder joint, a small number of grains comprised the entire

solder joint. While Figure 42b appears to show many small grains, it is more likely to be

interlaced dendrite arms from two to three grains.28

If cross-sectioned through another

plane, this solder joint would likely look similar to the other 3 alloys. The addition of Bi

to the solder does not appear to have a noticeable impact on the number of grains. Further

study is needed to understand the impact of Bi on the size and structure of Sn grains.

67

Figure 42: Polarized light images BGA a) SAC305 b) Senju M42

c) Sunrise and d) Sunflower

6.2 Comparison of Interfacial IMC Layers

IMC bond layer thicknesses are difficult to measure because of their uneven

morphologies; scalloped in the case of Cu6Sn5 and a needle shaped “coral” structure in

the case of (Cu,Ni)6Sn5. In order to determine a representative thickness, the following

method was used. For BGA solder joints, a cross-section was evaluated using the SEM.

Three solder joints were selected, one from each side and one from the middle of the

component. For each joint, three locations were selected, again one from each side and

one in the middle of the solder joint. For these nine locations, an image was recorded at

3000x magnification. This image was divided along the horizontal length into six equal

segments. Five measurements were then taken between the six segments. A total of 45

measurements were taken, representing various locations across the component. This

methodology allows for any difference in IMC thickness, due to thermal differentials

68

across a component, to be captured. Also, it ensures random locations are selected as

opposed to exclusively selecting the high or low points for measurement. This same

process was used for determining the IMC thickness of the QFP solder joints, however

only two leads were analyzed, resulting in a total of 30 measurements.

There is no statistically significant difference in the thickness measurements of the IMC

layers formed using each of the four alloys on the board side of the BGA. Each solder

paste formed a board side interfacial IMC layer with a normally distributed thickness of

approximately equal variance. The variation in thickness (Figure 43) is large, which can

be attributed to the scalloped shape of the (Cu,Ni)6Sn5 IMC bond layer, which forms at

the solder/board interface (Figure 45b). Table 20 and Table 21 provide a summary of test

for equal variance () and analysis of variance (ANOVA) testing of the means (µ) of the

various alloys for BGA and QFP solder joints respectively. These tables provide

probability (p-values) for each test. In all cases, the null hypothesis (H0) and alternate

hypotheses (Ha) are as follows:

H0: µSAC305=µSenju M42=µSunrise=µSunflower and SAC305=Senju M42=Sunrise=Sunflower

Ha: at least one µ is different and at least one is different

Values in Table 20 and Table 21 which are bold italicized represent points with a p-value

less than 0.05. This indicates that, within a 95% CI, the H0 is rejected and the Ha is

assumed to be valid.

In the case of the board side IMC layer of BGA solder joints, the H0 is valid, the IMC

layers are statistically similar.

In the case of the component side IMC bond layer, there is a statistical difference in the

thickness measurements: H0 is rejected due to the very low probability. SAC305 and

Sunrise appear to have a greater mean thickness than Senju M42 and Sunflower, however

there are many outliers within the distribution, corresponding to the random, long,

“needle-shaped” shapes which occur in the (Cu,Ni)6Sn5 phase (Figure 46). These points

were removed and the analysis was performed a second time. The new distribution was

69

also found to that the mean IMC thicknesses where not equal however this time, it was

only Sunflower exhibited a greater mean thickness. The new analysis also included many

new outliers. The jagged nature of these IMC layers makes it very difficult to compare

mean thicknesses of the component side IMC; the analysis was deemed to be

inconclusive.

Component SideBoard Side

SunflowerSunriseSenju M42SAC305SunflowerSunriseSenju M42SAC305

8

7

6

5

4

3

2

1

0

IMC

Th

ickn

ess

(µ

m)

Figure 43: IMC measurements of BGA solder joints after reflow (U204)


thickness of the BGA IMC layer

Variable / p value

Total IMC

Equal

variance ANOVA

Location

Board Side 0.36* 0.366

Component Side 0.09**

0.07**

0.005 (with outliers)

0.011 (without outliers)

*Bartlett’s Test for equal variance used since distributions where found to be Normal.

**Levene’s Test for equal variance used since distributions was found to be non-Normal.

70

The mean thicknesses of the IMC bond layer formed in the QFP solder joints are similar

for all four alloys in the “as manufactured” state (Figure 44). It should be noted that

SAC305 has a thicker IMC layer than the other three alloys at t = 0 if a 93% CI is used.

This is expected since the process temperature is higher than for the other three alloys.

Lead SideBoard Side


8

7

6

5

4

3

2

1

0

IMC

Th

ickn

ess

(µ

m)

Figure 44: IMC measurements of QFP solder joints after reflow (U2)


thickness of the QFP IMC layer

Variable / p value

Total IMC

Equal

variance ANOVA

Location

Board Side 0.60 0.07

Component Side 0.30 0.15

71

Figure 45: IMC bond layers formed on a) QFP components and b) BGA

components

The composition of the IMC bond layers was determined using EDX analysis. Although

EDX provides only a semi-quantitative analysis of the composition, it has been deemed

sufficient for the purpose of identifying phases. The Sn portion of this IMC layer is

contributed from the solder, all four solder alloys contain between 91.3 and 96.5 wt%

Sn. The Cu portion of the IMC is contributed from both the solder (all four solder alloys

contain between 0.5 to 0.75 wt% Cu), and the joining material (lead or board side Cu

pad).

Table 22: IMC type

Paste Alloy

BGA QFP

Board Side Component

Side Board Side Lead Side

SAC305 (Cu,Ni)6Sn5 Ni23Cu33Sn44,

(Cu,Ni)6Sn5 Cu6Sn5 Cu6Sn5

Senju M42 (Cu,Ni)6Sn5 Ni23Cu33Sn44,


Sunrise (Cu,Ni)6Sn5 Ni23Cu33Sn44,


Sunflower (Cu,Ni)6Sn5 Ni23Cu33Sn44,


The IMC bond layers which form on the BGA components are a combination of the

solder paste (one of the four tested alloys), SAC305 from the component solder ball, and

72

the material finish at both the component interface and the board side interface (Figure

45b). The component side interface consists of a Cu pad with a Ni finish

barrier/combined solder (paste alloy + ball alloy). On the board side, the interface

consists of a Cu pad/combined solder. EDX analysis shows that the IMC formed at the

component side is some combination of Ni23Cu33Sn44 and (Cu,Ni)6Sn5 (Figure 46). The

morphology of these needle-like, “coral” shaped IMCs accounts for the large spread in

thickness measurements (component side of Figure 43) and the number of large values

found outside of the 75% quartile denoted as “*” in Figure 43.

Figure 46: IMC formed on BGA at component side

spectrum 1) Ni23Cu33Sn44 and spectrum 2) (Cu,Ni)6Sn5

Figure 47: Typical morphology of a) Cu6Sn5 IMC as formed on a QFP solder joint

with Sunflower b) and c) (Cu,Ni)6Sn5 and Ni23Cu33Sn44 IMCs respectively both

formed on a BGA with SAC305 solder paste

73

7 Summary of Findings

Each of the four alloys under investigation produced solder joints of good shape and

acceptable degree of mixing on both QFPs and BGAs. The Bi content of the three Bi-

containing alloys allowed for melting temperatures of 14-18C lower than that of

SAC305 without negatively impacting the resulting solder joint formation. Differences in

microstructure have been observed related to the decrease in Ag content, addition of Bi

and varying amounts of Cu dissolution resulting from different process temperatures.

The QFP solder joints formed with Senju M42 were very similar to those formed with

SAC305, with all of the Bi present in the Sn phase. This dissolved Bi may act as a solid

solution strengthener, the effect of which needs to be verified in mechanical testing. Both

Sunflower and Sunrise showed some degree of Bi precipitating out of solid solution. Both

fine and medium sized particles were observed. This may achieve some precipitation

hardening within the solder joints, but again will need to be verified by further testing.

The final composition of BGAs was similar to that of SAC305, which was the main

contributing alloy to the final composition, and in all cases having a final Bi

concentration of 1wt% or less. All of the Bi was present in solution within the Sn

phase. The presence of many large, complex Cu6Sn5 flower-shaped imtermetallics

indicates some degree of undercooling in solder joints formed in the BGAs using

SAC305 as the ball material and each of the four alloys as the paste material.

The QFP interfacial IMCs at the board interface and lead interface were similar for all

four alloys. SAC305 results in a thicker IMC layer forming within a 93%CI. In all cases

Cu6Sn5 was identified. The BGA also showed no difference in the type of interfacial

IMCs that formed at both the board side and at the component side. There appears to be

no statistically significant difference in the layer thicknesses of the IMC formed on BGA,

although this is harder to evaluate with the Ni23Cu33Sn44 layer that forms at the

component side of the BGA due to the variable nature of the IMC shape.

74

8 References

1 A. Zbrzezny “Characterization and Modeling of Microstructural Evolution of Near-

Eutectic Sn-Ag-Cu Solder Joints” Ph.D. Thesis, Dept. MSE, Univ. Toronto,

Toronto, Canada, 2004.

2 L. Nie M. Osterman and M. Pecht “Copper Pad Dissolution and Microstructure

Analysis of Reworked Plastic Grid Array Assemblies”, at IPC APEX EXPO

Conference, Las Vegas, NV, 2009.

3 M. Kelly, C. Hamilton and P. Snugovsky “Have High Cu Dissolution Rates of

SAC305/405 Alloys Forced a Change in the Lead Free Alloy Used During PTH

Processes”, SMTA Proc. Pan Pacific Microelectronic Jan. 2007.

4 D.A. Porter and K.E.Easterling, “Thermodynamics and Phase Diagrams” in Phase



pp 48-55.

5 http://www.metallurgy.nist.gov/phase/solder/agcusn-ll.jpg

6 L.Snugovsky, P. Snugovsky, D. Perovic; T. Sack; J. W. Rutter “Some aspects of

nucleation and growth in Pb free Sn–Ag–Cu solder”, Mater. Sci. Technol. 21

(2005) 53-60.

7 A.R. Zbrzezny “Microstructure Characterization of Sn-Ag-Cu Lead-Free Solder

Solidified at Different Cooling Speeds”, Microsc. Microanal. 8 (Suppl. 2), 2002.

8 T-K. Lee, T. Bieler, C-U Kim, H. Ma “Phase Equilibria and Microstructure of Sn–Ag–

Cu Alloys” in Fundamentals of Lead-Free Solder Interconnect Technology, New

York, Springer, 2015 ch 3, pp. 64

9 Y Takamatsu, H Esaka, K Shinozuka “Formation Mechanism of Eutectic Cu6sn5 and

Ag3Sn after Growth of Primary –Sn in SN-Ag-Cu Alloy”, Materials

Transactions Vol. 52, No. 2, pp. 189-195, 2011.

10 J.W. Elmer, E.D. Specht, M. Kumar “Microstructure and In Situ Observations of

Undercooling for Nucleation of -Sn Relevant to Lead-Free Solder Alloys”,

Journal of Electronic Materials, Vol. 39, No.3, 2010.

11 P. Snugovsky, Z. Bagheri, C. Hamilton “Microstructure and Reliability Comparison of

Different Pb-Free Alloys Used for Wave Soldering and Rework”, Journal of

Electronic Materials, Vol. 38, No. 12, 2009.

12 T. Hurtony A Bonyár, P Gordon, G Harsányi “Investigation of intermetallic

compounds (IMCs) in electrochemically stripped solder joints with SEM”

Microelectronics Reliability 52, 2012.

75

13 T. Hurtony, A Bonyár, P Gordon, G Harsányi. “Investigation of intermetallic

compounds (IMCs) in electrochemically stripped solder joints with SEM”

Microelectronics Reliability 52, 2012.

14 A. So and Y.C. Chan “Reliability Studies of Surface Mount Solder Joints – Effect of

Cu-Sn Intermetallic Compounds” IEEE Transactions on Components, Packaging,

and Manufacturing Technology – Part B, Vol 19, No 3, August 1996.

15 Z Mei, A.J. Sunwoo, J.W. Morris “Analysis of Low-Temperature Intermetallic Growth

in Copper-Tin Diffusion Couples”. Metallurgical Transactions A , Volume 23A,

March 1992.

16 K.H. Prakash, T Sritharan “Interface Reaction Between Copper and Molten Tin-Lead

Solder”, Acta Mater. 49 (2001) 2481-2489.

17 S. Kumar, C.A. Handwerker and M.A. Dayananda “Intrinsic and Interdiffusion in Cu-

Sn Systems”, Journal of Phase Equilibria and Diffusion Vol. 32 No. 4, 2011.

18 K-N. Tu “Copper Tin Reactions in Bulk Samples” in Solder Joint Technology :

Materials, Properties and Reliability New York, Springer, 2007, ch. 2, sec. 2.6.1,

pp. 58-59

19 A.J. Sunwoo, J.W. Morris, G.K. Lucey, “The Growth of Cu-Sn Intermetallics at a

Pretinned Copper-Solder Interface”, Metallurgical Transactions A, Volume 23A,

April 1992

20 S. Kumar, J. Smetana, D. Love, J. Watkowski, R.Parker, C.A. Handwerker,

“Microvoid Formation at Electrodeposited Copper-Solder Interface During

Annealing: A Systematic Study of Root Cause”, presented at SMTAI, Chicago, Il,

2008.

21 H.F. Zou, H.J. Yang, Z.F. Zhang “A Study on the Orientation Relationship Between

the Scallop-Type Cu6Sn5 Grains and (011) Cu Substrate using Electron

Backscattered Diffraction” Journal of Applied Physics, 106, 113512 (2009)

22 K-N. Tu “Solder Reactions on Nickel, Palladium, and Gold” in Solder Joint

Technology: Materials, Properties and Reliability New York, Springer, 2007, ch.

7, pp. 183-204.

23 J-Y Tsai and J. Gaida “The Interaction between SnAgCu Solder and Ni(P)Au,

Ni(P)PdAu UBMS”

24 K Sweatman, J. Read, T. Nishimura and k. Nogita “The Effect of Microalloy Additions

on the Morphology and Growth of Interfacial Intermetallic in Low-Ag and No-Ag

Pb-Free Solders”, presented at SMTAI, Chicago, Il, 2011.

25 L. Snugovsky, P. Snugovsky, D.D. Perovic, J. W. Rutter “Phase Equilibria in Sn Rich

Corner of Cu-Ni-Sn System” Materials Science and Technology, 2006, Vol 22,

No 8. pp. 899-902

76





27 D.A. Porter and K.E.Easterling, “Diffusional Transformation in Solids” in Phase



pp 263-381.

28 B. Arfaei, N. Kim and E.J. Cotts “Dependence of Sn Grain Morphology of Sn-Ag-Cu

Solder on Solidification Temperature”, Journal of Electronic Materials, Vol. 41,

No. 2. 2012.

77

Chapter 3 Accelerated Thermal Cycling

1 Accelerated Testing for Reliability Analysis

Reliability can be defined as the probability of a product to meet specifications over a

given time period while subjected to determined environmental conditions.1 The

reliability of any electronic assembly depends on the reliability of each individual

element. An assembly is likely to fail due to component failure in the short term and due

to solder attach failure in the long term.2

The solder attach, which include the solder joint

and the material to which it adheres, particularly in the case of surface mount devices,

presents a complicated situation in that it acts as both the electrical contact as well as the

mechanical attachment.3 In field operation, a solder attach will experience loading

conditions in the form of: (i) mechanical load, (ii) vibration, (iii) thermal shock and (iv)

differential thermal expansion. The accumulated damage caused by one or more of these

stress conditions will eventually lead to wear-out failure. Therefore, determination of the

useful life of an electronic product is contingent upon the understanding of the failure rate

of the surface mount solder attach.

Accelerated testing is used to obtain more information than would be practical, or even

possible, under normal field conditions where years may pass before a sufficient number

of failures occur to determine the reliability of a product. Accelerated testing is also

essential in determining the effects of changes to product design through its design cycle.

Accelerated conditions may include a test environment in which conditions are more

severe than that experienced during normal equipment use or by increasing the frequency

in which a condition is applied to a product. In both cases, care should be taken to avoid

the introduction of failure mechanisms, which would not be encountered in field

operation of a product.4 If failure mechanisms are maintained, the accelerated data can be

used to extrapolate the expected failure rates in field conditions.2

Of the four loading conditions described above, thermomechanical fatigue resulting from

cyclic differential thermal expansion is the primary failure mechanism of surface mount

78

solder joints and the focus of this chapter.3,5

This form of fatigue is the combination of

plastic and creep deformation which occurs as a solder attach is exposed to cyclical

heating and cooling through powering on and off, as well as exposure to environmental

conditions. Accelerated Thermal Cycling (ATC) is used to test the response of a product

to rapidly repeating changes in temperature to a level higher than normally experienced

by a product in field conditions. In this case, the applied cyclical stress is in the form of

differential thermal expansion and contraction between the various materials within a

circuit board. These differences are defined by the coefficient of thermal expansion

(CTE) of the differing materials that make up the system. In a circuit board, there are two

levels of recognized differential thermal expansion:

1. Global: thermal expansion mismatch between components and substrate, as illustrated

in Figure 48.

2. Local: thermal expansion mismatch between solder and material to which it is

bonded6

Figure 48: Representation of stress generated in circuit due to CTE mismatch7

Stress-strain hysterisis loops, such as the one illustrated in Figure 49, are often used to

describe the cumulative stress history experienced by a solder attach during ATC. As the

temperature increases, the differential expansion of the various materials results in the

79

initial applied stress, which leads to plastic deformation. During the hot dwell (or hold)

period, which generates stresses lower than the yield strength of the solder, creep and

stress relaxation within the solder material occur. During accelerated testing, there may

or may not be enough time allotted for full stress relaxation as would likely exist in field

conditions. Stress is then applied in the opposite direction – the stress axis representing

an absolute value – during the cold cycle. As the cycles are repeated, the solder joint

fails due to low-cycle fatigue and can be described by the Coffin-Manson relationship

given in Equation 3-1:

m

pf CN1

)(1 1-1

where fN is the average number of cycles to failure, C1 and m are material constants.8

Greater temperature fluctuations or greater differences in CTE mismatch will result in

higher values of plastic shear strain ( p ). This value will also be impacted by

component geometry. For example leaded components will result in lower shear plastic

strain than leadless components. Ramp rate and dwell time, at the temperature extremes,

will also impact the shape of the hysteresis loop and the strain level. The Coffin-Manson

model assumes that plastic strain is the main deformation mechanism. This simplified

model has since been revised to include the more dominant creep effect, as well as

cyclical parameters. These models are presented in IPC-SM-785 and take into account

such factors as component and solder fillet geometry, potential cyclic fatigue damage at

complete stress relaxation, and fatigue ductility coefficients. These models are also

presented with numerous caveats indicating which conditions are required for each model

and what limitations may exist.2 Modeling of complex creep-fatigue in solder attach,

particularly for new solder alloys, is a continuing field of research.

80

Figure 49: Hysteresis loop for thermal cycle9

In order to relate the results of ATC testing, or any Accelerated Life Testing (ALT), to

life condition, data from at least two conditions need to be used to establish a trend line in

order to extrapolate to field condition. Low-level and a high-level acceleration condition

should be chosen. Three or more conditions would allow for determination of non-linear

relationships. Low-level acceleration should produce a mean time-to-failure (MTTF) of

about 10-20 times shorter than the actual field life. High-level acceleration should

produce a MTTF of approximately 100-500 times shorter than actual field life.2

Acceleration Factors (AF) are then calculated as the ratio between the two conditions and

then extrapolated from the chosen conditions to field conditions.

ATC qualification test conditions for various applications have been determined and

compiled in specification IPC-9701A6. These tests have been devised as minimum

criteria for accelerated testing for qualification purposes. While Appendix 1 of this

specification does provide guidance and examples for determining AF, it is not necessary

to determine these for every test. Rather, the minimum requirements per industry have

been established. This specification also states at least 32 samples per condition should

be tested and the test should run for the required duration or at least until 63.2% failure in

order to characterize the failure distribution. Tests that are stopped at the end of the

81

required Number of Thermal Cycle (NTC) level with insufficient (or no) failures will be

characterized by failure analysis (FA) alone. FA is to be performed on a minimum of

three randomly chosen samples. Table 23 provides a summary of the test conditions laid

out in this standard and Table 24 provides some examples of worst-case scenarios for

various products in field conditions.

Table 23: Sample of temperature cycling requirements Table 4-1 in IPC-9701A6

Test

Condition

Mandated

Condition Test Duration

Low

Temperature

Dwell

High

Temperature

Dwell

Temperature

Ramp Rate

TC1 0C to

100C

NTC-E

6,000 cycles

(preferred for

TC1) 10 minutes

(+0C/-5C)

10 minutes

(+5C/-0C)

20C/minute

TC4 -55C to

125C

NTC-C

1,000 cycles

(preferred for

TC4)

Table 24: Worst case use environments of SMT6

Application Tmin

(C)

Tmax

(C)

T

(C) Cycles/year

Typical service

life (years)

Acceptable

failure risk (%)

Consumer 0 60 35 365 1-3 1

Military Aircraft

a)

b)

c)

-55 125

40

60

80

100

100

65

10-20 0.01

82

Note in Table 24, T is not simply the difference between Tmax and Tmin representing the

absolute maximum and minimum temperatures experienced in the given environment,

but rather the expected worse-case scenarios which occur in a given service condition.2

Solder in ATC testing has been shown to fail according to a creep-fatigue model resulting

from the applied thermomechanical fatigue stresses. The mechanical properties of solder

are temperature dependent and therefore change continuously over the course of a fatigue

cycle, both in the bulk microstructure of the solder as well as in the IMC layer formed

between the solder and the Cu substrate. There are also cumulative effects of repeated

cycles.10

The total strain (T) can be expressed by Equation 3-2 where e and p are the

elastic and plastic strains respectively and c represents creep strain.

cpeT 1-2

During the high temperature dwell period of the thermal cycle, creep and stress relaxation

are the dominant strain evolution mechanisms. During ramp up and down, plastic and

elastic strain deformations are dominant unless this transition is slow enough to allow for

full stress relaxation, which is not typically the case in accelerated testing. It is the

repeated plastic and elastic strain which constitutes fatigue fracture.

2 Microstructural Evolution

2.1 Changes to Bulk Solder

Notably, many of the tests and models described above have been developed for SnPb

solders, which have been tested over decades. It has been shown that the microstructural

response to creep and fatigue behavior of Sn based SAC alloys is very different than that

of SnPb.11

Therefore, it is possible the models and even the applicable test conditions

need to be updated. In SnPb solder joints, the initial response to applied

thermomechanical stress is grain coarsening in an attempt to reduce the internal energy in

the smaller grains. Both the applied shear strain which results from CTE mismatch and

the elevated temperature aid in grain coarsening. In eutectic SnPb, the Sn rich regions and

83

Pb regions will coalesce into larger islands. During coarsening, it should be noted that

the IMC attach layer also thickens – a Cu3Sn region may form between the Cu substrate

and the Cu6Sn5. Immediately adjacent to the IMC a layer of Pb-rich phase will form as

the local Sn has been consumed in the IMC growth process. Pb also tends to pool around

the grain boundaries of the Sn rich phase. Grain coarsening is accomplished via grain

boundary sliding and grain boundary diffusion-induced migration.5

Microvoids begin to

appear along grain boundaries, interphase boundaries and interfaces with IMCs.

Microvoids continue to propagate along grain boundaries or other interfaces until they

connect to form microcracks and eventually catastrophic macro level cracks. Figure 50

illustrates a typical SnPb solder attach a) before any accelerated testing and b) after 3000

cycles of harsh ATC. These images, taken at the same magnification, illustrate the

degree of grain coarsening, IMC growth and phase coalescence that occurs during

repeated stress cycles. While this joint did not fail catastrophically after 3000 cycles, a

large crack is propagating through the bulk solder along a Pb-rich region and, if testing

had continued, would have lead to a full electrical and mechanical fail.

Figure 50: SnPb solder a) before testing and b) after 3000 cycles -55C to 125C12

The microstructure of solder attach formed with SAC type alloys are found to behave

differently from that of eutectic SnPb. As described in Chapter Chapter 2, these solders

are made up of up to 98% Sn with small amounts of Ag and Cu present as intermetallics.

The Sn matrix, as well as the Ag3Sn and Cu6Sn5 precipitates all influence the creep

response of the solder attach. As the strain increases, an initial coarsening of the IMC

particles occurs. Recrystallization of the Sn grains takes place in the area of high strain,

84

usually close to the bulk solder interface with the attach IMCs. The grain boundaries

formed between these new grains, as well as some large precipitates, provide sites along

which voids can form (at the triple points) and fractures can easily propagate.10,13

Additionally, it is during the hot dwell cycle that the brittle IMC attach layer may grow.7

Figure 51 shows a solder attach made up of SAC305 a) before testing and b) after 3000

cycles of harsh testing. From Figure 51a) we can see the dendritic arms of the Sn

appear relatively small and evenly spaced with small dispersiods of Ag3Sn in the

interdendritic spacing. After ATC, the Ag3Sn particles appear to have coalesced into

fewer, larger particles. Additionally, the IMC layers formed between the lead and solder

as well as between the solder and copper pad have increased in size and appear to have

two phases, Cu3Sn and Cu6Sn5 as seen in Figure 51b). A Cu3Sn layer was not detectable

at zero time. Finally, the dendritic structure does not appear to be well defined after

thermal cycling. Polarized light or electron backscatter diffraction (EBSD) is needed to

properly view the grain structure. Figure 52 shows an example of a BGA solder joint that

failed during thermal cycling; recrystallized grains appear in the region where high shear

strength is experienced, and where the final crack propagated. Using EBSD, it has been

proposed that this recrystallization occurs ahead of the propagating crack.10

Fractures can

easily propagate along the regions created under creep conditions, notably between grain

boundaries and along the bulk solder/IMC interface.

Figure 51: SAC305 solder a) before testing and b) after 3000 cycles -55C to 125C12

85

Figure 52: SAC305 after ATC shown with a)polarized light and b)EBSD mapping 13

2.2 Changes to Interfacial IMC during Accelerated Thermal Cycling

A good interfacial IMC layer is required for a properly formed solder joint, as it provides

the mechanical, electrical and thermal connection between the bulk solder and the Cu

substrate. This layer, however tends to be brittle and if a very thick layer forms or

develops, it may become a reliability concern. Interface layers also tend to be preferential

sites for crack formation and propagation. Additionally, the IMC layer in solder joints

may be a reliability concern because it consumes Cu from the substrate, Cu being the

dominant diffusing species.14

The consumption of Cu may result in a weakened Cu pad

as described in 2.1.

Cu3Sn was not detected upon initial solidification (Chapter Chapter 2); it forms as a result

of solid-state reaction during thermal aging at temperatures above 60C.17

The Cu3Sn

growth occurs between the Cu6Sn5 and the Cu substrate and is governed by the

availability of Cu from the substrate and Sn either from the already present Cu6Sn5 as

described in equation 3-3 or Sn from the bulk solder diffusing through the Cu6Sn5

substrate as described in equation 3-4. The initial formation of Cu6Sn5, and its continued

growth are dependent on Sn from the bulk solder and Cu from the substrate. The Cu is

either present during the initial solidification reaction, or diffuses through the Cu3Sn

layer. Both are described by equation 3-5.15

86

Cu6Sn5 + 9Cu 5Cu3Sn 2-1

3Cu + Sn Cu3Sn 2-2

6Cu + 5Sn Cu6Sn5 2-3

The Cu3Sn layer is of concern for three reasons: it is a brittle layer more so than Cu6Sn5

or bulk solder, it is associated with large volume shrinkage, and the growth of a Cu3Sn

layer next to Cu is often accompanied by Kirkendall voids along the interface of the two

phases.16

These small voids may form a weak interface.

Both Ag and Cu additives to the solder have been found to be beneficial in suppressing

the growth of the Cu3Sn layer.17

Ag does not participate in the interfacial IMC layer at all

and can therefore only act as an influence by retarding the diffusivity of Sn through the

bulk solder towards the IMC layer. Other alloying elements (e.g. Ni, Zn) have also been

found to retard the growth of Cu3Sn and in some instances Cu6Sn5. In these cases, trace

amounts of the element are found in, or at the interface of the IMC layer and therefore are

thought to participate directly in the reaction or provide a diffusion barrier between the

layers.18

The change in thickness of the intermetallic layers can be described by the Arrhenius

equation (3-6) where x is the thickness, x0 is the initial thickness in m, t is time in s, T is

temperature in K, R is the universal gas constant (8.314 kJ/mol K) and A, Q and n are

material constants. 20

)exp(0 RTQ

Atxx n 2-4

2.3 Effects of Bi

Bi as an alloying element is expected to have an overall influence on the reliability of a

solder joint in two ways. The first is the reduced process temperature due to the addition

of Bi, which may result in a reduced thickness of the interfacial IMC layer as well as

87

reducing some of the deleterious effects of high temperature on both board material and

components.19

The second involves modification of the bulk solder microstructure. Sn phase is

considered to have most potential for modification with alloying as it makes up most of

the volume within a SAC solder joint. Alloys with Bi have been explored as a means to

improve thermomechanical properties through solid solution strengthening and

precipitate hardening.20

In a previous study, for example, after 3000 cycles harsh

environment ATC (-55C to 125C), Bi which was previously present in solid solution

with Sn precipitated out in an evenly distributed manner, reducing microstructural

degradation.12

Other studies have shown that Bi segregates at the grain boundaries,

potentially reducing creep resistance.21

The influence of Bi on the formation of interfacial IMC layers has also been explored. It

has been shown that Bi does not participate in the formation of either intermetallic

compound layers (Cu6Sn5 or Cu3Sn), but there is a higher concentration within the Sn

close to the Cu6Sn5 suggesting that Bi is rejected from the Sn as Sn is contributed to the

formation of Cu6Sn5.22

In this thesis, the behavior of Bi containing alloys is investigated. Also, the impact of

reducing the amount of Ag in SAC-based solders is also examined.

3 Experimental Setup

The work in this section forms a screening experiment in which the performance of

selected alloys was evaluated during thermal cycling under two different conditions. The

intention was to provide data for determining the suitability of these solders for more

statistically relevant reliability testing, which is too costly to perform on all combinations

of solder, board material and temperature conditions. These tests were performed in

accordance with test conditions outlined in IPC-9701A6 and IPC-SM-785

2. The statistical

requirements of these specifications were not adhered to due to limited resources.

88

ATC testing assumes that solder joints were properly formed and no wetting issues exist,

which would result in infant mortality due to defective solder joints.6

The quality of

solder joints used in this study was demonstrated to be satisfactory in Chapter Chapter 2.

In addition to testing a statistically relevant number of samples, IPC-9701A states that the

duration of testing would ideally generate at least 63.2% failures within that sample set in

order to properly evaluate reliability and possibly determine acceleration factors. This

specification states that if an insufficient number of failures, or no failures at all, are

generated at the end of the number of thermal cycles (NTC) level, then failure analysis

should be performed on a random selection of parts (a minimum of three per test

variation) to verify that no failure occurred. This failure analysis can also be useful in

determining whether or not any microstructural changes occurred during the testing.

Finally, as this testing is intended to address solder joint reliability, only those failures

determined to be the result of thermomechanical component/board interaction will be

considered. Board level failures such as via cracks, delamination, will not be included in

the data set for any solder joint reliability calculations.6

3.1 Materials

Three Bi-containing alloys, Senju M42 containing 3%Bi and Sunrise and Sunflower each

containing 7wt% Bi, with low or no Ag, were tested against SAC305. Two board

materials were tested; a normal Tg material– one which is typically used for SnPb

applications around 150C, and a high Tg material– one which is now required for most

Pb-free applications at around 170C. Table 25 summarizes the combinations of solder

and board material tested. As described in 2.2, the lower Tg board materials that had been

previously used with SnPb solders had a lower elastic modulus and were less susceptible

to pad cratering, delamination and warpage failures.

89

Table 25: Build matrix for ATC testing

Alloy Composition Board Material Assembly Temperature

SAC305 Sn 3%Ag 0.5%Cu

High Tg

(170C)

240C





Normal Tg

(150C)

240C




3.2 Test Vehicle

Celestica’s RIA3 test vehicle, shown in Figure 53, was selected as the test vehicle. It was

designed to simulate a typical, medium complexity SMT assembly. It is an 8”x10”

surface made up of 12 copper layers for a total thickness of 0.093” with an Organic

Solderability Preservative (OSP) finish. This board is often used to test new, lead-free

solders and other material parameters. LQFP176, PBGA256, CBGA64 and MLF20s were

populated, two of each per board. The BGA components all had SAC305 ball alloy and

all component groups were reflowed in an air environment. Figure 53 and Table 26

shows the test vehicle and highlights the monitored components.

Figure 53: Test vehicle with monitored components

90

Table 26: Monitored components for ATC testing

Component Reference Designator

1206 1206-A1

1206 1206-B1

0402 R400-R419

0402 R420-R439

CSP64 U208

CSP64 U209

MLF20 U5

MLF20 U6

PBGA256 U204

PBGA256 U205

QFP176 U1

QFP176 U2

3.3 Test Strategy

All components were monitored using electrical data loggers; failures were defined as a

20% increase in nominal resistance over five consecutive scans. Failures were cut from

the boards, and then the rest of the components were returned to the chamber for further

testing. Each of the two thermal profiles was allowed to run for a preset number of

cycles.

Figure 54: Card set up in thermal cycling chamber

91

3.3.1 0-100C Thermal Cycling

This portion of the testing was done in accordance with IPC-9701A Test Condition 1

(TC1) of 0C (+0/-5C) to 100C (+5/-0C) for Number of Thermal Cycles (NTC-E),

which is preferred for TC1, of 6000 cycles. The boards under test, listed in Table 27 were

set up in a chamber using a racking system which allowed for airflow around the boards

(Figure 54). The final thermal cycle profile is shown in Figure 55.

Table 27: Test matrix for 0C to 100C ATC

Alloy Composition Board Material # of Boards Tested

SAC305 Sn 3%Ag 0.5%Cu High Tg (170C) 4

Normal Tg (150C) 3


3%Bi

High Tg (170C) 6

Normal Tg (150C) 5

Sunrise Sn 1%Ag 0.7%Cu 7%Bi High Tg (170C) 5

Normal Tg (150C) 6

Sunflower Sn 0.7%Cu 7%Bi High Tg (170C) 5

Normal Tg (150C) 5

Figure 55: Chamber profile for 0C to 100C thermal cycling

92

3.3.2 Harsh Environment (-55C to 125C) Thermal Cycling

This portion of testing was done in accordance with IPC-9701A Test Condition 1 (TC4)

of -55C (+0/-5C) to 125C (+5/-0C) for NTC-C of 1000 cycles. Notably, the

maximum temperature selected for the thermal cycle profile should be 25C lower than

the Tg of the board material.6 In this case the Normal Tg board material is 150C, while it

is exactly 25C lower than the Tg, does not allow for the 10C leeway on the hot end of

the cycle making the board material vulnerable to damage during testing. The boards

under test are listed in Table 28. The final thermal cycle profile is shown in Figure 56.

Table 28: Test matrix for -50-125C ATC

Alloy Composition Board Material # of Boards Tested

SAC305 Sn 3%Ag 0.5%Cu High Tg (170C) 3

Normal Tg (150C) 3

Senju M42 Sn 2%Ag 0.75%Cu 3%Bi High Tg (170C) 5

Normal Tg (150C) 5

Sunrise Sn 1%Ag 0.7%Cu 7%Bi High Tg (170C) 5

Normal Tg (150C) 6

Sunflower Sn 0.7%Cu 7%Bi High Tg (170C) 5

Normal Tg (150C) 5

93

Figure 56: Profile for -55C to 125C thermal cycling

3.4 Post ATC Evaluation and Failure Analysis

Samples from groups exposed to each of the two ATC profiles were evaluated to

determine both the failure modes and to examine the changing microstructure over the

course of the test. Samples were selected at time of failure, at the halfway point of each

test and at the completion of each test. Each sample was mounted in epoxy, and then

ground and polished through the following sequence: 500 and 1200 grade SiC paper,

polishing with 6 µm and 1 µm DiaPro diamond suspensions (Struers), and an oxide

polish (Struers’ OP-S). Optical microscopy was performed using a Nikon Measurescope

MM-11. Prior to SEM analysis, the samples were carbon coated using an Emitech

K950X. SEM microscopy was performed using a Hitachi S-4500 and Hitachi S-3000N

with Oxford and ThermoScientific EDX systems respectively.

94

4 Results

4.1 Reliability and Failure Analysis Results

Using IPC-9701A as a guideline, this test focuses on solder joint reliability; other modes

of failure, for example those attributed to the laminate material, were removed from

consideration in the reliability calculations. Therefore any identified failures that were

removed from the test prior to completion and later found to be attributed to something

other than solder joint fatigue should be identified as “censored”. Finally, testing was

suspended after a predetermined amount of time whether or not failures were identified.

Parts that did not fail were identified as “censored”. This type of data censoring is

referred to as Type 1, or time censored data in which a test is concluded after a

predetermined amount of time, in spite of the fact that failures may not have occurred.23

4.1.1 0C to 100C Accelerated Thermal Cycling

A total of 6010 cycles of ATC from 0-100C were completed. The test was periodically

stopped and the failures, which have been recorded using a data logger, were removed

and verified by manual resistance measurement. If the failure was determined to be

within the test vehicle, as opposed to a cabling issue, the component was cut out of the

board for future failure analysis. The remaining components on the test vehicle where

returned to the chamber and the testing continued. Because no failures had been

identified at the halfway point of the test, four boards, one for each alloy, were removed

for metallurgical analysis. This occurred after 3157 cycles had been completed. These

components were identified as “censored”.

Table 29 summarizes the number of cycles to first failure of the QFP176s monitored

during the 6010 cycles of the test. There are not enough failures at the end of 6010

cycles to plot failure distribution charts for each alloy. Only SAC305 experienced failures

when built on High Tg (170C) board material and those occurred towards the end of the

test. Figure 57 shows the distribution of failures of solder joints built on High Tg boards

95

compared with those built on Normal Tg boards. A two-parameter Weibull distribution is

used.

The shape parameter determines the shape of the probability distribution and can be

viewed by plotting the probability density function (PDF). At values between 3 and 4, the

shape parameter approximates a normal distribution. Shape parameter values higher than

4, as shown in Figure 58 results in a near-normal distribution with a left tail. The scale

parameter (or slope on the Weibull distribution) described the range of the distribution, or

the rate of at which the failure rate increases.

It can be seen that the High Tg boards are expected to survive much longer than those

built with Normal Tg when using SAC305 as the solder paste.

Table 29: Summary of QFP failures after 0C to 100C ATC

Component

type

High Tg Normal Tg

SAC305 Senju

M42 Sunrise Sunflower SAC305

Senju

M42 Sunrise Sunflower

QFP176 2/6 0/10 0/7 0/8 6/6 0/10 1/12 0/10

33.3% -- -- -- 100% -- 8.3% --

Failure

Cycle

5792 3534 5416

5994 4349

4653

4782

5223

5251

96

Figure 57: Weibull plots of SAC305 QFP solder joints after 6010 Cycles 0 to 100C

comparing High Tg to Normal Tg boards

Figure 58: Probability Density Function for SAC305 QFP solder joints after 6010

Cycles 0 to 100C comparing High Tg to Normal Tg boards

Figure 59 shows the failure occurred as a result of fracture initiating in the bulk solder

and propagating along the lead side IMC layer. This is further illustrated in Figure 60

which shows the fracture surface of a similar solder joint made up of SAC305 on High Tg

97

board material. The area highlighted in red shows the crack initiation. The area

highlighted in green illustrates the point at which the fracture moves along the IMC layer.

Within the green area, EDX scans were taken of area 1, which was shown to be pure Sn

indicating bulk solder and area 2: which was shown to be Cu6Sn5 indicating the IMC

layer. Finally, the area highlighted in blue shows a ductile fracture, which likely occurred

as a result of overload on the remaining area.

Figure 59: QFP fracture of SAC305 on High Tg board a) optically and b) cross

section

98

Figure 60: Fracture surface of QFP solder joint with SAC305 on High Tg board

after 6010 cycles

Figure 61 shows that fractures had initiated in solder joints formed with the other three

alloys on high Tg board material, but by 6010 cycles had not yet propagated far enough

into the solder joint to cause an electrical failure. It is expected that these solder joints

would ultimately fail in the same manner as shown in Figure 59 and Figure 60 if the

testing had continued.

99

Figure 61: Fracture initiation in QFP a) Sunflower b) Senju M42 after 6010 cycles

As described in 6.1.2, the BGA solder joints were made up of paste from the

experimental alloy and SAC305 solder balls. This resulted in a solder joint composition

closer to SAC305 and with no more than 0.9wt% Bi. No failures occurred within the

6010 cycles of the test when high Tg board material was used. A number of failures

occurred for each of the paste alloys when normal Tg board material was used as outlined

in Table 30. Using resistance probing, these were all found to be the result of failure

within the component-solder-board system; either in the board circuit pattern beneath the

component, the solder connecting the component to the board or within the component

itself. It is not possible to determine where within this system the failure occurred

without destructive techniques. Each component was therefore cut from the board and

cross-sectioned to find the point of failure.

Table 30: Summary of BGA failures after 0C to 100C ATC

Component type Normal Tg

SAC305 Senju M42 Sunrise Sunflower

Fraction of components failed 5/6 9/10 8/12 6/10

% of components failed 83.3% 90.0% 66.7% 60.0%

Cycles to Failure

4168 4079 4534 3623

4612 4540 4619 3930

5168 4629 4725 4607

5168 4680 4855 4902

5278 4861 5153 5086

5497 5452 5169

5674 5455

5679 5536

5871

100

Figure 62 illustrates that there is no significant difference in the reliability performance of

the solder joints formed using paste of the four alloys on Normal Tg boards. All BGA

failures were found to result from board failures, specifically via cracks within the

laminate material. The probability distributions for each alloy in this test exhibit similar

shape and scale parameters, which is to be expected since all failures are characteristics

of the board material and construction, rather than the solder alloy.

Figure 62: Weibull plots of BGA solder joints on Normal Tg boards after 6010 Cycles 0 to

100C cycles comparing Four Alloys

Figure 63 illustrates a) a typical via barrel failure and b) a bulk solder fracture – although

incomplete failure – near the component side IMC.

101

Figure 63: BGA failures after 0 to 100C ATC a) failure in board material by via

plating crack and b) partial failure through bulk solder (SAC305) near component

side IMC

4.1.2 -55C to 125C Accelerated Thermal Cycling

A total of 1000 cycles of ATC from -55C to 125C, or Harsh Environment ATC, were

completed. The test was periodically stopped and the failures, recorded using a data

logger, were removed and verified by manual resistance measurement. If, during the

course of testing, the failure was determined to be within the test vehicle, as opposed to a

cabling issue, the component was cut out of the board for future failure analysis. The

remaining components on the test vehicle were returned to the chamber and the testing

continued. Four boards, one for each alloy, were removed for metallurgical analysis near

the halfway point of the test, after 438 cycles completed. Table 31 summarizes the

number of cycles to first failure of the QFP176s monitored during the 1000 cycles of the

test.

102

Table 31: Summary of QFP failures after -55C to 125C ATC

Component

type

High Tg Normal Tg

SAC305 Senju


Senju


Fraction of

components failed 3/4 6/10 4/9 4/10 4/5 6/8 10/12 5/10

% of components

failed 75.0% 60.0% 44.4% 40.0% 80.0%

75.0

% 83.3% 50.0%

Cycles to

Failure

457 406 81 493 304 428 83 23

514 446 212 689 544 439 83 480

820 485 526 897 721 561 285 662

535 595 909 918 637 459 697

621 643 528 783

634 680 601

726

744

843

924

Figure 64 and Figure 65 show probability distribution plots and the probability density

functions of the four alloys on High Tg board materials respectively. Sunrise showed

some early fails, resulting in a different shape of the PDF.

Figure 66 shows probability distribution plots of the four alloys on Normal Tg board

materials respectively. In all cases, the distribution of failures occurred within

overlapping confidence intervals; there is insufficient data to distinguish between the

alloys after 1000 cycles of testing on Normal Tg boards.

103

Figure 64: Weibull plots of QFP failures on High Tg boards after 1000 cycles -55 -

125C

Figure 65: Probability Density Function for QFP failures on High Tg boards after

1000 cycles -55 -125C

104

Figure 66: Weibull plots of QFP failures on Normal Tg boards after 1000 Cycles -55-

125C

Just as in the 0 to 100C cycling test, all QFPs experienced solder fracture initiating in

the bulk solder and propagating along the lead side IMC (Figure 67). These however did

not result in complete fracture, and would not account for the electrical failure. Further

testing, by probing various test points in the test vehicle, determined that the likely cause

of electrical failure was the copper trace separating from the board material. Copper trace

fractures result in electrical open circuits.

105

Figure 67: QFP176 fractures in a) Sunflower and b) Senju M42 after 1000 cycles

Harsh testing

The BGA components that underwent harsh thermal cycling experienced many failures

(Table 32). Upon cross sectional failure analysis however, these failures were found to be

the result of board damage, as in the 0-100C test, specifically via cracks within the

laminate material (Figure 68). While there are enough data points to plot probability

distributions, this did not provide a means of distinguishing between alloys.

Table 32: Summary of BGA failures after -55C to 125C ATC

Component type High Tg Normal Tg

SAC305 Senju


Senju


Fraction of

components failed 5/6 10/10 6/6 9/10 4/6 10/10 12/12 10/10

% of components

failed 83.3% 100.0% 100.0% 90.0% 66.7% 100.0% 100.0% 100.0%

Cycles to Failure

407 702 376 522 498 387 325 425

449 757 518 575 581 459 425 469

455 772 537 670 615 461 443 561

815 783 757 692 655 464 447 600

847 783 783 693 541 467 643

798 823 696 552 471 672

878 864 570 491 705

924 968 597 494 804

985 999 600 575 837

999 696 587 934

848

897

106

Figure 68: BGA failures after -55 to 125C ATC failures in board material by via

plating crack in a) boards built with Sunrise and b) boards built with Sunflower

Further examination of the solder joints revealed some crack initiation in the solder joints

towards the component side of the BGA solder joint. These crack initiation sites (Figure

69) were less significant than the damage observed after 6010 cycles of 0-100C thermal

cycling and would not account for a failure reading. Electrical failures, just as in the case

of 0-100C thermal cycling, occurred as a result of via cracks in the board material

(Figure 68). A probability plot comparing all failures on the two different board materials

(Figure 70) illustrates that the board material was a more significant distinguishing factor

than the various solder alloys.

Figure 69: BGA solder crack initiation after -55 to 125C ATC in solder joints

formed with SAC305 solder balls and a) Sunflower and b) Sunrise

107

Figure 70: Weibull plots of BGA solder joints after 1000 Cycles -55 to 125C cycles

comparing two board materials

QFP solder joint failures on Normal Tg board material was used to compare the two test

conditions, 0-100C for 6010 cycles and -55 to 125C for 1000 cycles. The Harsh

environment test (-55 to 125C) produced many failures over the 1000 cycles, including a

few early failures. This resulted in a Weibull distribution with a shape parameter of 1.26,

a PDF with a right tail. The 0-100C was less severe and therefore resulted in fewer

failures and no early failures. The shape parameter of 5.59 indicates that the failure rate

increases over time.

108

Figure 71: Weibull plot for QFP solder joints on Normal Tg boards comparing two

test conditions

Figure 72: Probability Density Function for QFP failures on Normal Tg boards

comparing two test conditions

109

4.2 Microstructure Evaluation

4.2.1 Bulk Microstructure

SAC305, before thermal cycling (Figure 73a and 73b) appears as small dendrite arms

surrounded by small Ag3Sn and Cu6Sn5 particles. After thermal cycling from 0 to 100C

for 3148 and 6010 cycles (Figure 73c and 73d respectively) the intermetallic particles

coalesced into fewer, larger particles. Senju M42 behaved in the same manner. In both

cases, the bulk of the microstructural transition occurred between 0 and 3148 cycles, with

little change between 3148 and 6010 cycles. With 3% Bi in Senju M42, there does not

appear to be any significant Bi precipitation from the solid solution (Figure 74).

Figure 73: SAC305 at Time 0 a) optically and b) SEM image and after b) 3148 cycles

and d) 6010 cycles of 0-100C ATC

110

Figure 74: Senju M42 at 1000x after a) 3148 cycles and b) 6010 cycles

Sunrise and Sunflower solder paste alloys, which both contain 7wt% Bi, showed a

significant amount of Bi precipitation after ATC. In both cases large, uneven particles of

Bi were present upon solidification, as seen in Chapter 2. After ATC, Sunrise showed

many, very small particles precipitated throughout the bulk, evenly dispersed from the

Sn grains (Figure 75). Sunflower also exhibited some Bi accumulation along the grain

boundaries (Figure 76). In both cases, there did not appear to be any significant

accumulation of Bi along the interface with the interfacial IMC layer.

Figure 75: Sunrise after 6010 cycles 0-100C at a) 500x and b) 1000x

111

Figure 76: Sunflower at a) Time 0 and b) after 6010 cycles 0-100C at 500x

4.2.2 IMC Growth during Thermal Cycling

The interfacial IMC layers were measured as described in section 6.2. A comparison of

the board side IMC layer of a QFP176 component follows. This IMC layer formed

during the reflow process in which a bond was formed between the board side copper pad

and the molten solder. The Time 0 results reference this IMC layer after reflow but

before any thermal cycling. As described in section 6.2, only Cu6Sn5 was identified at

Time 0. Measurements were made after approximately the halfway point and again at the

completion of the two thermal cycling regimes. For the 0-100C test, measurements were

completed after 3157 cycles and then again after 6010 cycles. For the harsh environment

test (-55C to 125C) measurements were taken at 438 cycles and then again at 1000

cycles. After exposure to thermal cycling, both Cu6Sn5 and Cu3Sn species were identified

(Figure 77).

112

Figure 77: IMC layers formed on QFP176 solder joints between the board side Cu

layer and solder paste a) sunrise and b) SAC305 after 438 cycles of Harsh thermal

cycling. Location 1 shows the Cu3Sn layer, location 2 shows the Cu6Sn5 layer

Figure 78 provides a comparison of the overall IMC thickness at the various test

intervals; Figure 79 provides a comparison of the Cu3Sn layer within the IMC at various

test intervals. Table 33 and Table 34 provide a summary of results from two-sided T-test

to compare the mean values and Levine-test to compare the variance of the IMC and

Cu3Sn thicknesses respectively. These tables are provided in order to examine the

changes in thickness of the overall IMC and the Cu3Sn layers, for each alloy, over the

course of the two thermal cycling conditions. The tables provide probability (p-values)

for each test. In all cases, the null and alternate hypotheses are as follows:

H0: µ1=µ2 and 1=2

Ha: µ1≠µ2 and 1≠2

Values in Table 33 and Table 34 which are bold italicized represent points with a p-value

less than 0.05. This indicates that H0 should be rejected at a 95% confidence level and

the Ha is assumed to be valid. In this case, it assumes that the mean value has changed.

These results show that between 0 and 1000 cycles of Harsh testing, the IMC layers for

each alloy increased significantly. For Senju M42 and Sunrise, this occurred primarily

within the first 438 cycles. For SAC305 and Sunflower the increase continued throughout

113

the test. In 0-100C ATC testing, the IMC layer for all alloys except SAC305 increased

throughout the test. The IMC thickness for SAC305 increased most significantly during

the first half of the test.

All t-tests of the µ in Table 34 result in rejecting the H0. This indicates that the thickness

of the Cu3Sn layer continues to increase throughout the course of the two tests.

Table 35 provides a summary of test for equal variance and ANOVA testing of the

various alloys. This test compares the IMC and Cu3Sn thicknesses of all four alloys, at a

given point during ATC. This table provides probability for each test. In all cases, the

null and alternate hypotheses are as follows:



The result of ANOVA testing, indicated in Table 35: Results of ANOVA test to equal

variance and compare means, show that while the total IMC thickness increased

relatively consistently among the four alloys, the increase in the thickness of the Cu3Sn

layer was different in at least one alloy. Figure 79 clearly shows that Sunflower exhibited

a significantly thicker Cu3Sn layer after Time 0 in both test conditions.

114

Figure 78: IMC growth at the board side of QFP during thermal cycling

Figure 79: Cu3Sn (portion of IMC) growth at board side of QFP during thermal

cycling

115

Table 33: Results of Levine-Test to compare the variance () of IMC measurement

and 2-sided t-Test to compare the means (µ) of IMC measurements

Test

Condition

HARSH

(-55 to 125C) (0 to 100C)

Comparison 0 to 438

cycles

0 to 1000

cycles

438 to 1000

cycles

0 to 3157

cycles

0 to 6010

cycles

3157 to

6010 cycles

Variable /

p value µ µ µ µ µ µ

SAC305 0.40 0.00 0.21 0.00 0.04 0.03 0.23 0.00 0.11 0.00 0.74 0.56

Senju M42 0.72 0.00 0.17 0.00 0.11 0.49 0.53 0.00 0.97 0.00 0.61 0.04

Sunrise 0.03 0.00 0.13 0.00 0.43 0.85 0.53 0.00 0.02 0.00 0.40 0.00

Sunflower 0.28 0.00 0.75 0.00 0.26 0.00 0.31 0.00 0.87 0.00 0.39 0.00

Table 34: Results of Levine-Test to compare the variance () of Cu3Sn measurement

and 2-sided t-Test to compare the means (µ) of Cu3Sn measurements

Test

Condition

HARSH

(-55 to 125C) (0 to 100C)

Comparison 438 to 1000 cycles 3157 to 6010 cycles

variable µ µ

SAC305 0.40 0.00 0.72 0.00

Senju M42 0.70 0.00 0.99 0.00

Sunrise 0.17 0.04 0.17 0.00

Sunflower 0.01 0.00 0.04 0.00

116

Table 35: Results of ANOVA test to equal variance and compare means

Variable / p value

Total IMC Cu3Sn

Equal

variance

()

ANOVA

(µ)

Equal

variance

()

ANOVA

(µ)

Test

Condition

Time 0 0.60 0.07 -- --

438 cycles

HARSH (-55 to 125C) 0.20 0.04 0.01 0.00

1000 cycles

HARSH (-55 to 125C) 0.35 0.24 0.00 0.00

3157 cycles

(0 to 100C) 0.65 0.30 0.56 0.00

6010 cycles

(0 to 100C) 0.13 0.43 0.01 0.00

Figure 80 through Figure 87 provide details of the ANOVA analysis. The interval plot

Figure 80 and Figure 81 show the 95% confidence intervals for the mean overall IMC

layer thickness and the mean Cu3Sn layer thickness respectively. While it appears that

SAC305 has a thicker overall IMC layer at Time 0, the interval plot shows substantial

overlap between the 95% CIs of all four alloys and therefore indicates that the difference

in thickness is not statistically significant. Further analysis indicates that there is a

difference in Time 0 thicknesses within a 93% confidence level. The thicker IMC layer

of SAC305 is a result of the higher processing temperature (240C). The intervals for the

Cu3Sn layer thickness of Sunflower do not overlap with any of the 95% CI of the other

alloys, thereby indicating a significant difference in thickness.

117

Figure 80: Interval plot of IMC thickness at the board side of QFP after ATC

Figure 81: Interval plot of Cu3Sn thickness at board side of QFP after ATC

118

The main effects plots in Figure 82 to Figure 85 are used to de-couple the main factors in

the measurements to better understand their individual contributions. Figure 82 and

Figure 83 show the condition, or time at exposure to thermal cycling, has a greater

influence on the overall IMC thickness than do the particular alloys. SAC305 has a Time

0 IMC thickness greater than the mean of the other three alloys, which all have mean

values closer to the grand mean.

Figure 84 and Figure 85 show that the condition is also a greater influence on the mean

overall thickness of the mean Cu3Sn layer than the alloys. The difference between Cu3Sn

thicknesses amongst the alloys shows the same pattern: SAC305 and Senju M42 have

mean thicknesses close to the grand mean, Sunrise has a mean lower than the grand mean

and Sunflower is significantly greater than the grand mean.

Figure 82: Main effects plot of IMC thickness at the board side QFP during HARSH

(-55 to 125C) ATC

119

Figure 83: Main effects plot of IMC thickness at the board side QFP during 0-100C

Figure 84: Main effects plot of Cu3Sn thickness at board side of QFP during

HARSH (-55 to 125C) ATC

120

Figure 85: Main effects plot of Cu3Sn thickness board side of QFP during 0-100C

Interaction plots found in Figure 86 and Figure 87 tests the factors to identify any

interactions. Parallel, or near parallel lines, as seen in both cases indicate there are no

clear interactions; the influence of one factor is not dependent on the other. The mean

thickness of the IMC layer for SAC305 at Time 0 is greater than that of the other alloys,

however over the course of ACT testing under two test conditions, this distinction seems

to disappear, the mean IMC thicknesses are no longer distinguishable between alloys.

121

Figure 86: Interaction plot of IMC thickness at the board side of QFP after ATC

Figure 87: Interaction Plot of Cu3Sn Thickness at Board Side of QFP after ATC

122

5 Summary of Findings and Conclusions

5.1 Findings Based on Reliability Data

During two different thermal cycling regimes the only failures which could be attributed

to solder joint fatigue failures occurred in SAC305 during 0 to 100C cycling; one

sample of Sunrise, which failed but only after 5416 cycles. All other electrical failures

were attributed to failures within the board material. In the case of BGAs, this occurred

due to cracks within via barrels; in QFPs it was the result of copper trace fractures.

Thermal cycling did not significantly distinguish between the reliability performances of

the four alloys. High Tg board material appeared to outperform Normal Tg board material.

It can be concluded that the three lower process temperature solders performed as well as,

or better than SAC305 during accelerate thermal cycling. All solders survived longer

than the board materials.

5.2 Findings Based on Microstructural Observations

Bi had a beneficial impact on the bulk microstructure. It precipitated from the bulk solder

evenly in very small particles. This favorable influence of Bi only appeared when the

concentration was 7wt%; the 3wt% Bi of Senju M42 did not appear to have the same

impact. Senju M42 appeared to have the same aging characteristics as SAC305. There

was some degree of Bi segregation along the grain boundary observed in the Sunflower

alloy; however there did not appear to be full segregation of the phases.

The interfacial IMC layer () Cu6Sn5, formed during solidification, was similar in

thickness for all alloys. During thermal cycling, the total thickness of the interfacial IMC

continued to increase, as well as () Cu3Sn layer formation. While the overall thickness

increased similarly for all four alloys, the Cu3Sn layer of the Sunflower alloy was much

thicker than the other alloys. This alloy has no Ag, which indicates that Ag plays a role

in suppressing Cu3Sn growth. This is in spite of the fact that Ag does not participate in

the IMC formation; no Ag-containing species exists at the interfacial IMC. It has

previously been shown that eutectic SnAg solder exhibited a lower layer-growth

123

coefficient for Cu6Sn5 than for Cu3Sn.24

This was believed to be the result of the Sn

diffusion in SnAg being slower than the Sn diffusion in, for example eutectic SnPb

solder, in turn favoring the growth of Cu3Sn. It appears, through this study, that even

small amounts of Ag in the solder (1%, 2% and 3%) are sufficient to suppress the

diffusion of Sn to the growing interfacial IMC layers. Bi alloying did not significantly

change the growth rates of either intermetallic compound. The increased growth of the

Cu3Sn phase in the Sunflower alloy did not correspond to any decrease in reliability

during ATC.

SAC305, which was processed at a higher temperature, initially had a larger interfacial

IMC layer. This effect however did not continue during thermal cycling where the IMC

layers of all four alloys formed and increased to similar overall thicknesses.

124

6 References

1 J. Bentley “Introduction to Reliability and Quality Engineering”, 2

nd Ed, Essex,

England, Pearson Education Limited, 1999, ch. 2, pp 28-43.

2 IPC-SM-785 Guidelines for Accelerated Reliability Testing of Surface Mount Solder

Attachments, November 1992

3 W. Engelmaier “Solder Attachment Reliability, Accelerated Testing, and Results

Evaluation” in Solder Joint Reliability J.H. Lau, Ed. New York, Springer 1991,

ch.17. pp. 545-587

4 Blueprints for Product Reliability, The Reliability Information Analysis Center (RIAC),

http://theriac.org/DeskReference/viewDocument.php?id=280&Scope=blueprints

&Deskref=blueprint1#3point4

5 J.W.Evans “Thermomechanical Fatigue” in A Guide to Lead-free Solder: Physical

Metallurgy and Reliability, W. Engelmaier, Ed. London, UK, Springer 2007, ch.7,

pp. 145-185

6 IPC-9701A: Performance Test Methods and Qualification Requirements for Surface

Mount Solder Attachments, February 2006

7 A. MacDiarmid “Thermal Cycling Failure – Part One of Two” in The Journal of the

Reliability Information Analysis Center, January 2011

8 J-P.M. Clech and J.A. Augis “Surface Mount Attachment Reliability and Figures of

Merit for Design for Reliability in Solder Joint Reliability J.H. Lau, Ed. New

York, Springer 1991, ch.18. pp. 588-613

9 J.W.Evans “Introduction to Solder Alloys and Their Properties” in A Guide to Lead-free

Solder: Physical Metallurgy and Reliability, Ed. London, UK, Springer 2007,

ch.1, pp. 1-27

10 A. Zbrzezny “Characterization and Modeling of Microstructural Evolution of Near-

Eutectic Sn-Ag-Cu Solder Joints” Ph.D. Thesis, Dept. MSE, Univ. Toronto,

Toronto, Canada, 2004.

11 R. Coyle, R. Parker, M. Osterman, S. Longgood, K. Sweatman, E. Benedetto, A. Allen,

E. George, J. Smetana, K. Howell, J. Arnold “iNemi Pb-Free Alloy

Characterization Project Report: Part V – The Effect of Dwell Time on Thermal

Fatigue Reliability” presented at SMTAI, Chicago, Il, 2013




Vegas, NV, 2014.

125

13 B. Arfaei, M. Anselm, S. Joshi, S. Mahin-Shirazi, P. Borgesen, E. Cotts, J. Wilcox, and

R. Coyle “Effect of Sn Grain Morphology on Failure Mechanism and Reliability

of Lead-Free Solder Joints in Thermal Cycling Tests”, presented at SMTAI,

Chicago, Il, 2013

14 K-N. Tu “Copper Tin Reactions in Thin-Film Samples” in Solder Joint Technology :

Materials, Properties and Reliability New York, Springer, 2007, ch. 3, pp. 73-

108.

15 N. Mookam and K. Kanlayasiri “Evolution of Intermetallic Compounds between Sn-

0.3Ag-0.7Cu Low-Ag Lead-free Solder and Cu Substrate during Thermal Aging”

J. Mater. Sci. Technol., 2012, 28(1), 53-59.

16 C. Yu, Y.Yang, P. Li, J. Chen, H. Lu “Suppression of Cu3Sn and Kirkendall voids at

Cu/Sn-3.5Ag solder joints by adding a small amount of Ge”, J.Mater Sci: Mater

Electron (2012) 23:56-60

17 C. Yu Chen, Kai-Yun Wang, Jing-Qing Chen, Hao Lu “Suppression effect of Cu and

Ag on Cu3Sn layer in solder joints”, J Mater Sci: Mater Electron (2013) 24:4690-

4635

18 G.C. Moon, S.K. Kang, D-Y Shin and H.M. Lee “Effects of Minor Additions of Zn on

Interfacial Reactions of Sn-Ag-Cu and Sn-Cu Solders with Various Cu Substrates

during Thermal Aging” Journal of Electronic Materials, Vol. 36, no. 11, 2007.





20 P. Vianco, and J.A. Rejent “Properties of Ternary Sn-Ag-Bi Solder Alloys: Part 1 –



21 D.Witkin “Creep Behavior of Bi-Containing Lead-Free Solder Alloys” Journal of

Electronic Materials, Vol. 41, No. 2, 2012.

22 P. Vianco, and J.A. Rejent “Properties of Ternary Sn-Ag-Bi Solder Alloys: Part 1 –



23 CRE Primer, Chapter IX: Data Collection, p. IX-5, Quality Council of Indiana. 2009.

24 D.Kwon “Intermetallic Formation and Growth” in A Guide to Lead-free Solders:

Physical Metallurgy and Reliability, Silver Spring, MD: Springer, 2005, pp. 97-

126

126

Chapter 4 Tin Whisker Testing

1 Introduction

Whiskers, spontaneous columnar or cylindrical filaments which emanate from a surface,

are found to form from various metals including Ag, Zn, Cd and Sn.1,2

The danger posed

within an electronic system is that whiskers are both conductive and grow spontaneously.

It is therefore challenging to predict where they could form, and if they would grow long

enough to cause potential electrical bridging issues. Electrical bridges can form by

whiskers growing long enough to cross the minimum gap, or by whiskers broken off

from the surface and falling freely around the electronic assembly. Furthermore, there is a

potential for two whiskers to grow sufficiently close together to allow for electrical

arcing between them. As the demand for increasingly small electrical components

continues, and the minimum lead-to-lead gap shrinks, the risk whiskers pose increases

with miniaturization.

Sn whiskers tend to have the following characteristics: growth over time, thicknesses

varying from sub-micron to just above a few microns, tough and electrically conductive.3

Although Sn whiskers were first observed shortly after WWII in telephone transmission

line channel filters1, they were not extensively studied since Pb in Sn (at least 1wt%) was

found to be an effective form of mitigation. Although whiskers can grow from SnPb

solder, they were observed as short and often topped with a Pb-rich “cap” (Figure 88),

which appeared to limit their growth.

127

Figure 88: Pb "cap" on whisker from SnPb component finish2

Initial investigations of whisker formation focused on Sn plating of electronic

components. As whiskers are a surface, rather than a bulk, phenomenon, a thin coating of

matte Sn found on most Pb-free components, was initially thought to be the main area of

concern. Correspondingly, most standards developed for testing, measuring and

qualifying new materials focused on component surface finishes.4,5

During assembly using a SnPb solder, the solder is expected to wet a portion of the lead

and form a fillet (represented by E in Figure 89a) while the rest of the lead maintains its

original plating material. Due to the higher reflow temperatures used to assemble SAC

solders, the electroplated Sn on the lead frame melts. The solder is therefore likely to

completely wet and cover the lead (Figure 89b).6 With SnPb solder assemblies, using

matte Sn plated components; the risk of whisker formation exists on the exposed Sn area

of the lead. Originally it was thought that whiskers grow primarily from thin film

coatings, however it has been shown they also grow from bulk solder; whiskers have

been observed growing from bulk SAC305 and SAC405.7 This indicates the risk of

whisker growth now spans the entire solder joint surface on a Pb-free assembly. It is

believed however, that areas where a thin layer of solder exists are of greatest concern.

128

Figure 89: a) Schematic for a typical solder joint of a leaded component using SnPb

solder8, b) cross section showing solder joint formed with Pb-free solder

1.1 Whisker Growth Kinetics

Currently there is no widely accepted explanation for the mechanism causing Sn

whiskers; however it is widely agreed that the driving force for whisker nucleation and

growth involves local compressive stresses. It is accepted that Sn whiskers are one of the

forms of stress relaxation within the localized region.6 It is also understood that whiskers

are a surface stress relief mechanism; when relaxation occurs within the bulk

microstructure, whiskers will not grow.9 Therefore areas with thin solder coating, or thin

plating material appear to be more susceptible to whisker formation. For example, on an

assembly of Sn plated components built with SAC305 solder, whiskers grew from areas

where the solder was less than 25µm thick.10

Complicating matters, it has been shown that the total length of the whisker does not

correlate to the applied stress, thereby indicating the mechanism is not one of bulk

diffusion. Additionally, whiskers appear to originate from newly formed grains at the

surface. For these reasons, one of the currently accepted models, proposed by Vianco,

describes whisker growth as a process of cyclic dynamic recrystallization (DRX).2,11

This model suggests that whiskers grow under specific conditions requiring both a cyclic

DRX, in order to nucleate a new grain at the surface of a thin film, and a mass transport

129

mechanism to facilitate the supply of Sn. In both case the driving force is applied

compressive stress. This model also proposes that stress relaxation will occur by means

of recrystallization rather than recovery when the homologous temperature (Th) is

approximately 0.6. Sn, even at room temperature, has a homologous temperature of 0.59.

The formation of whisker growth is further dependent on strain rate where a slower strain

rate will favor whisker formation over a high strain rate. Finally if the applied stress is

too high, even at a low strain rate, whiskers will not likely form.

As compressive stress is applied to a surface, strain within the localized system

accumulates, usually in the form of dislocations at pre-existing grain boundaries. Once

this strain energy pile-up exceeds a certain limit, a new grain will nucleate and grow in a

process known as DRX. This process differs from a dynamic recovery process in which

the dislocations will either annihilate each other or be absorbed within the grain

boundaries instead of forming new grains. The growth of grains in the form of whiskers

represents a surface phenomenon rather than grain growth within the bulk of a material.

Strain rate, temperature and grain size all factor into the cyclic DRX model and

subsequently whisker growth.2

Figure 90: Cyclic Dynamic Recrystallization resulting in whisker formation12

Creep Deformation

Compressive Stress

Dynamic Stress Relaxation

Cyclic DRX Continuous DRX

– no whiskers

Work Hardening

Recovery

– no whiskers Recrystallization

(DRX)

X < σ < 5 MPa σ > 5 MPa

– no whiskers

– no whiskers

Th= ~0.6 Th= 0.2-0.3

Slow strain rate Fast strain rate

130

1.2 Sources of Compressive Stress

Sn whisker growth is initiated by self generating compressive stress within the material

and can be mechanical, thermal or chemical in nature.9 An IMC layer growing between

the lead frame material and the plating material, or solder, is usually irregular, or scallop

shaped (Chapter 3). The growth of this layer and any IMC particles, which subsequently

break off into the bulk material, introduce compressive stresses into the bulk material. In

the case of Sn plated leads, the Sn layer is often found to have a columnar shape, with

IMC particles accumulating along the grain boundaries.

Figure 91: Source of compressive stress contributing to whisker growth13

When exposed to various temperature cycling environments, stress induced into the Sn

plating or solder is attributed to the mismatch in the Coefficient of Thermal Expansion

(CTE) between the Sn and the lead frame material. Localized compressive stresses may

occur and increase the whisker propensity where thin layers of solder exist. Also, the

uneven interfacial IMCs, which continue to grow during the heating portion of the

thermal cycle, may introduce compressive stresses into the solder. There may not be

sufficient time to relieve stress before the cold portion of the cycle occurs. Previous work

found that CTE mismatch and corresponding whisker growth was greater for Alloy 42

lead frame materials than for Cu.10,13

(Table 36).

131

Table 36: CTE values for common materials in solder joints

Material Composition Coefficient of Thermal Expansion

(CTE) ppm/C

Lead Finish Sn Sn 22.0

Solder Paste

Materials

SnPb Sn - 37Pb 21.6

SAC305 Sn - 3Ag0.5Cu 21.6

Lead Frame

Materials

Alloy 42 Fe - 42Ni 5.8

Cu194 Cu - 2.1-2.6Fe-0.015-

0.15P-0.05-0.2Zn 17.5

Cu151 Cu - 0.1Zr 17.7

Cu7025 Cu - 2.2-4.2Ni0.25-

1.2Si0.05-0.3Mg 17.3

The formation of oxides on the surface of a solder joint or plated lead has also been

associated with whisker growth. 9,14

In this case, the compressive stresses are related to

the volume expansion caused by the formation of oxides, which can lead to a volume

expansion of approximately 29 to 34% as observed in many studies testing samples in

humid conditions. 15

Oxidation can also be associated with the presence of IMC particles.

For example, when exposed to humidity and chloride contamination, Ag3Sn IMC

particles found at the surface of a SAC305 solder joint are found to exhibit corrosion

within the interdendritic spaces (Figure 92). This is attributed to the differing chemical

potentials of the IMC and the surrounding bulk solder, which facilitate galvanic

corrosion.6,16

Sn-based solders are particularly susceptible due to the roughness of the

surface; shrinkage voids resulting from the dendritic solidification of these solders leaves

a very rough surface which can easily entrap contaminates during the manufacturing

process or during field use. Cu-based lead frame materials have been shown to be more

susceptible to these types of compressive stresses than Alloy 42.16

132

Figure 92: Ag3Sn oxide zone with whisker6

Mechanically-induced residual stresses also result in whisker formation. For example, the

compressive stresses generated in the Sn plating of a leaded component have been

attributed to the bending or stretching processes used to form leads into their final shape

after plating.17

Studies show that whiskers growing as a result of these mechanically

induced stresses tend to have a higher growth rate than those resulting from room

temperature IMC internal stresses.18

1.3 Morphology of Sn Whiskers

As stated in 1.1, a correlation between applied stress and resulting whisker length has not

been found. Whisker growth is, to date, unpredictable as is the final morphology of the

whisker. Long, thin, straight whiskers are often found alongside short, kinked whiskers of

various thicknesses (Figure 93). The long, fine whiskers have been found to contain only

Sn, usually of a single crystal structure. Shorter, thicker whiskers, hillocks and shell

shaped protrusions are made up of mainly Sn but may contain Ag3Sn and Cu6Sn5 IMC

particles.7

It is the long, thin, straight whiskers which are of most concern as they have

133

the highest potential of bridging a gap between electrical contacts. However, it is not

currently possible to predict the form or total length of whiskers.

Figure 93: Whisker morphology a) long, thin whiskers and b) short, kinked

whiskers16

1.4 The Effects of Bi in Solder on Whisker Formation

SnPb solder and plating finish has long been used in electronics without major reliability

concerns related to Sn whiskers. Eutectic SnPb plating over a Cu surface forms an

equiaxed grain structure rather than the columnar grain structure typical of Sn plating

(Figure 94). This equiaxed structure is believed to allow stress relaxation to occur by

incorporating displaced Sn grains more readily at grain boundaries, which lie parallel to

the surface. This would allow for the material to relieve the compressive stresses in a

uniform creep.19,20

Jadhav et al.19

showed that both the grain structure and the

concentration of Bi had an affect on the overall stress relaxation of a thin film over Cu.

Figure 94: Cross sections of plating surface made with SEM FIB of a) Sn and b)

SnPb19

134

Small amounts of Pb, as small as 1wt%, result in grain refinement and subsequently the

suppression of whisker growth. It was found that hillocks grow instead and that these

hillocks are of limited length.18

As Pb is no longer available as an alloying agent in

solder, other alloying elements are being explored for the same grain refinement, creep

resistance and ultimately whisker suppression characteristics. In this work, the focus is on

Bi alloying.

Section 1.2 identified the growth of an uneven, scalloped interfacial IMC layer as one of

the sources of compressive stress within a solder joint, particularly in areas with thin

solder coverage. Jo found that the IMC layer, which continued to grow during room

temperature storage and high temperature high humidity conditions, became more

uniform (i.e. less scalloped) with increasing concentrations of Bi.18

Annealing has also

been proposed as a means of creating a more uniform IMC layer. If the IMC layer were

to become more uniform in shape, it would introduce less compressive stress into the

system, thereby eliminating one possible source of stress. The surface area of the IMC

would also be reduced, which in turn would reduce the diffusion rate of Cu into the bulk

solder. The Cu6Sn5 IMC particles within the bulk solder are another source of internal

compressive stress on the localized system. Additionally, Jo found that as little as 0.5wt%

Bi reduced whisker propensity. Other researchers have indicated that as much as 3-5wt%

Bi is required for whisker suppression.1,21

The three main mechanisms by which Bi is thought to mitigate against whisker growth

are:

Refining the grain size

Altering the grain structure from a typical Sn columnar structure to an equiaxed

structure more similar to that found in SnPb

Reducing growth and irregularity of interfacial IMC18

135


The work describes in this section formed a screening experiment in which the selected

alloys were screened for whisker mitigation properties. The results were intended to

determine which, if any, of these alloys would be good candidates for further testing.

Table 37 provides a list of test conditions recommended by JEDEC for qualifying plating

finishes on components. It should be noted that there is no current test or qualification

guideline for assessing the propensity of solder alloys in an assembly to grow whiskers.

Further, it has been found that suspension of some tests, particularly those involving

humidity storage, will impact the total whisker growth and lead to under-reporting of

whisker length. When tests are restarted, new whiskers nucleate and grow rather than

existing whiskers continuing to grow. Therefore, in order to understand the time

dependence of whisker length, each successive inspection interval needs to be made after

an increased amount of uninterrupted exposure.6

The guidelines set out in Table 37 were

used as the basis for developing this works test conditions, however the conditions were

also augmented by further study. In high temperature/humidity storage conditions for

example, it was found that whisker nucleation occurred faster at 85C than at 50-70C.22

Table 37: JESD22A121.01 test conditions4

Stress Type Ref.

Spec. Test Conditions

Recommendations

Inspection

Interval

Minimum

Duration

Temperature Cycling JESD22

-A104

Min Temperature

-55 to -40 (+0/-10)C

Max Temperature

+85 (+10/-10)C

air to air: 5 to 10 minute soak;

~3 cycles/hour

500 cycles 1000 cycles

Ambient

Temperature/Humidity

Storage

302C and 603%RH 1000 hours 3000 hours

High

Temperature/Humidity

Storage

602C and 87 +3/-2%RH 1000 hours 3000 hours

136

2.1 Materials

Two alloys with low Ag (Senju M42 and Sunrise), and one alloy with no Ag (Sunflower),

and varying amounts of Bi, were tested against SAC305 (Table 38). The alloy selection

method was described in Chapter Chapter 1. QFP components (U1) were cut from the

test vehicle described in 5.1.

Table 38: Alloys screened for whisker growth


SAC305 Sn 3%Ag 0.5%Cu 240C




2.2 High Temperature High Humidity

Two samples of each alloy, consisting of 176 leads each, were placed in a humidity

chamber. Teflon cabling was used to suspend the samples (Figure 95) and de-ionized

(DI) water was used to generate the humidity conditions so as not to introduce possible

contaminates into the system. The samples were exposed to 85C/85%RH for 1000

hours.

Figure 95: Samples in HTHH chamber

137

2.3 Thermal Shock

Two samples of each alloy, consisting of 176 leads each, were placed in a two stage, air-

to-air, thermal shock chamber (Figure 96). The samples were exposed to -55C to 85C

thermal shocks using the profile shown in Figure 97 for a total of 1610 shock cycles.

Figure 96: Two stage, air to air, chamber for thermal shock testing

138

-80

-60

-40

-20

0

20

40

60

80

100

120

8:09

:04

8:11

:04

8:13

:04

8:15

:04

8:17

:04

8:19

:04

8:21

:04

8:23

:04

8:25

:04

8:27

:04

8:29

:04

8:31

:04

8:33

:04

8:35

:04

8:37

:04

8:39

:04

8:41

:04

8:43

:04

8:45

:04

8:47

:04

8:49

:04

8:51

:04

8:53

:04

8:55

:04

8:57

:04

8:59

:04

9:01

:04

9:03

:04

9:05

:04

9:07

:04

9:09

:04

9:11

:05

9:13

:05

Tem

pera

ture

in

°C

Cold Chamber Hot Chamber Sample

Figure 97: Thermal shock temperature profile

2.4 Post Exposure Evaluation

After each of the two exposures, components were inspected for whiskers using a

variable pressure SEM (Hitachi S-3000N) at 15 kV acceleration voltage (Vacc), and

25MPa vacuum pressure. All leads on the components where examined at low

magnifications, between 100x and 250x. Areas which showed signs of irregularity,

nucleation and/or whiskers growth, where then further examined at higher magnifications

of 1000 to 5000x.

The samples exposed to thermal shock, which subsequently showed whisker growth,

were then further evaluated by cross section. Each sample was first mounted in epoxy,

and then ground and polished through the following sequence: 500 and 1200 grade SiC

paper, polishing with 6 µm and 1 µm DiaPro diamond suspensions (Struers), and an

oxide polish (Struers’ OP-S). Optical microscopy was performed using a Nikon

Measurescope MM-11. Prior to SEM analysis, the samples were carbon coated using an

139

Emitech K950X. SEM microscopy on cross-sectioned samples was performed using a

Hitachi S-4500 and Hitachi S-3000N with the following EDX systems: Oxford and

ThermoScientific respectively.

3 Results

3.1 High Temperature High Humidity Results

No whiskers were found on any of the four alloys tested after 1000 hours exposure to

85C/85% RH. Further exposure, for at least an additional 3000 hours in the same

conditions is recommended to distinguish between the whisker growth propensities

amongst the test alloys. This would be closer to the conditions set out in

JESD22A121.01, however due to resource constraints was not completed as part of this

work.

3.2 Thermal Shock Results

The results of whisker inspection are summarized in Table 39 including the location of

whiskers, as described in Figure 98 and a summary of the Bi content found in the location

of whisker growth. All alloys were found to form whiskers under the thermal shock

conditions. All whiskers (or hillocks) where short, less than 10µm and would therefore

not be considered to fail JEDEC Standard No. 201A for class 2 components, which

requires that no whisker to exceed 45µm after 1000 cycles of thermal shock.5 While this

specification is intended for component finishes rather than soldered components, it is the

only currently available guideline by which to make a comparison.

140

Table 39: Summary of whisker growth after 1610 cycles thermal shock

Alloy Composition Whisker Location | Morphology wt% Bi at

Location

SAC305 Sn 3%Ag 0.5%Cu Yes 1,4 few hillocks 0

Senju

M42

Sn 2%Ag 0.75%Cu

3%Bi Yes

1,4 few, hillocks

1.3, 1.8 1,4

many, very small,

whisker nucleation sites

Sunrise Sn 1%Ag 0.7%Cu

7%Bi Yes

2 few, hillocks 3.7

1,4 many, very small

whisker nucleation sites 3.3, 3.3

Sunflower Sn 0.7%Cu 7%Bi Yes

1,2,4 few, hillocks 2.9, 4.1, 3.3

4 some, very small

whisker nucleation sites 3.3

Figure 98: Schematic showing locations on lead where whiskers formed

141

Along with the observed whisker growth, other competing stress relaxation mechanisms

were found in all solder alloys. Figure 99 shows two adjacent leads, each with very

different surface morphologies after 1610 thermal cycles. The lead on the left shows

massive, bulk deformation and volume recrystallization. The lead on the right exhibits

very little deformation and appears unaffected by the local stresses. Massive eruptions

and protrusions of bulk IMC particles were also observed.

Figure 99: Two adjacent leads with Sunflower solder paste after 1610 thermal

shocks

SAC305 solder joints exhibited a small amount of whisker growth after 1610 cycles

(Figure 100), however there were no whiskers longer than 10µm. Short, thick whiskers –

which can be classified as hillocks, were found in locations 1 and 4 (Figure 98). Location

1 is of greater concern as it has a short bridging distance with the adjacent lead. Both

locations are known to have thin solder coverage; a thin solder layer is more susceptible

to whisker growth as compared to an area with a larger volume of bulk solder which may

relieve internal stresses by competing mechanisms. Unlike the Bi containing alloys,

SAC305 did not exhibit any area with many, small leads or “fields” of very short and thin

whiskers.

All three of the Bi containing alloys exhibited some degree of hillocks growing amongst

fields of very short, very thin whisker nucleation sites. This was most often found in

locations 1 and 4 (Figure 101 to Figure 103). Sunflower exhibited the least area of these

142

fields but had the two single longest whiskers seen in this study. Both however were still

shorter than 10µm.

Figure 100: Whisker growth on SAC305 after 1610 thermal shocks. a) and b)

whisker growth in location 4 c) massive deformation and d) whisker growth in

location1

143

Figure 101: Whisker growth on Senju M42 after 1610 thermal shocks. a) and b)

location 4 with short, thick whisker surrounded by many, very short whisker

nucleation sites c) and d) short, thick whiskers growing at site of contamination in

location4

144

Figure 102: Whisker growth on Sunrise after 1610 thermal shocks. a) and b)

location 2 with short, thick whisker c) location 4 and d) location 1 with many, very

short whisker nucleation sites short

145

Figure 103: Whisker growth on Sunflower after 1610 thermal shocks. a) and b)

location 1 with short, thick whisker c) some, very short whisker nucleation sites at

location 4 and d) longest whisker observed at location1

Figure 104 shows a whisker growing from Sunflower solder after 1610 cycles of thermal

shock. The base of the new whisker grain is located at the intersection of three surface

grains. This finding aligns with the dynamic recrystallization model. A cross section of a

hillock formed from Sunrise (Figure 105), shows that the hillock forms from a new grain

or grains. Figure 106 shows mass bulk recrystallization of the Senju M42 alloy in

location 2, where the volume of solder is greater and more susceptible to bulk

deformation. Whiskers, being a surface phenomena, will not likely form in these regions.

The stress relaxation will likely occur as a result of one of the competing mechanisms.

146

Figure 104: Whisker growing from Sunflower at grain boundary of recrystallized

grains

Figure 105: Hillock growing from Sunrise after thermal shock

147

Figure 106: Senju M42 bulk recrystallization

Interfacial IMC layer growth has been identified as a potential source of compressive

stresses (section 1.2). The IMC may continue to grow during testing, particularly during

the high temperature portion of the cycle. If this growth continues in an irregular, scallop

shape, it will continue to drive compressive stress into the solder.

Figure 107 provides a comparison of IMC thicknesses after exposure. The IMC layers,

which form and grow at the board side of the solder joint, do not show significant

distinction between the solders. Table 40 provides a summary of test for equal variance

and ANOVA testing of the various alloys. This table provides probability (p-values) for

each test. In all cases, the null (H0) and alternate hypotheses (Ha) are as follows:



There is no statistically significant difference in the thickness measurements of the IMC

layers formed using each of the four alloys on the board side of the QFP after 1610 cycles

of Thermal Shock. However there is a difference in the thickness of the layer that at the

component side of the solder joint after Thermal Shock. In the case of the board side IMC

148

layer, the H0 is assumed to be valid while in the case of the component side, it is rejected.

The difference in variance of measurements for each alloy is not significant as

determined by a p-value 0.05. The difference in the mean (µ) thickness measurement at

the component side, in which SAC305 and Senju M42 have similar distributions, Sunrise

and Sunflower have similar, and smaller thickness distributions (Figure 108). This

suggests that the Bi content has some impact on the growth of the IMC layer under

thermal shock conditions. At the lead side, the mean of the IMC layer height is lower in

Sunrise and Sunflower than in SAC305 and Senju M42, which have roughly the same

overall height. The lead side IMC is of more interest than the board side as the whisker

growth has been observed mainly in locations 1, 3 and 4 (Figure 98) where a thin layer of

solder and the IMC layer interact potentially creating ideal conditions for whisker

growth. The board side IMC interacts with a larger volume of solder, at location 2. This

allows for more bulk recrystallization and massive deformation and may not drive the

growth of whiskers.

Lead SideBoard Side


8

7

6

5

4

3

2

1

0

IMC

Th

ickn

ess

(µ

m)

Figure 107: IMC measurements of QFP solder joints after 1610 cycles thermal

shock (U1)

149


thickness of the QFP IMC layer after 1610 cycles thermal shock.

Variable / p value

Total IMC

Equal

variance ANOVA

Location

Board Side 0.597 0.771

Lead Side 0.477 0.007

SunriseSunflowerSenju M42SAC305

2.50

2.25

2.00

1.75

1.50

IMC

Th

ickn

ess

(µ

m)

95% CI for the Mean - Lead Side

The pooled standard deviation was used to calculate the intervals.

Figure 108: Interval plot of IMC thickness at the lead side of QFP after thermal

shock

150

Figure 109: IMC layer at lead with a) SAC305 and b) Sunrise after 1610 cycles of

thermal shock

Figure 110 shows a whisker growing from SAC305 at location 4. The image on the right

shows that a thin layer of solder, approximately 10µm where the relative influence of the

IMC layer, which ranges from 1-4µm, may be great.

Figure 110: Whisker growing from SAC305 after thermal shock

An examination of the solder surface as well as cross sections revealed that the Bi is not

uniformly distributed through the solder joint. Figure 111 shows the Bi concentration at

various locations along the solder joint formed using Sunrise solder paste after 1610

cycles of thermal shock. The Bi content is lowest in the thin areas of the lead, those most

susceptible to whisker formation.

151

Figure 111: Bi content at various locations of a Sunrise solder joint

In both Sunrise and Sunflower, where the concentration of Bi is 7wt%, the Bi

precipitates, seen as the lighter areas in Figure 112, have been found around the grain

boundaries. Bi is also seen precipitating out of the primary Sn dendritic structure and

can be seen as fine particles, particularly visible in the top two images of Figure 112.

152

Figure 112: Cross section of Sunflower showing Bi accumulating at grain

boundaries

4 Summary of Findings and Conclusions

The whisker testing executed comprised a screening experiment. A systemic count and

measuring scheme, similar to those carried out in other studies10,13,16

is required to

provide statistical conclusions. In this study, two samples of each alloy, consisting of

176 leads each were tested in each environmental condition. This provides a large

enough sample size, however it was found that the length of test, specifically the HTHH

produced an insufficient number of whiskers and no whiskers of excessive length, which

would be of concern. The current required length of test is 3000 hours; in this

experiment, only 1000 hours was completed. From previous work6, it is known that

interruption of test corresponds to the interruption of whisker growth. It is therefore

recommended that the samples be reexamined after an additional 3000 hours of HTHH.

1610 cycles of Thermal Shock testing is sufficient to meet the requirements for whisker

testing. In this case, all four alloys passed the requirements set out in JEDEC Standard

153

No. 201A for class 2 components. It is believed that the internal, compressive stresses

induced by this type of thermal shock are very high (Figure 90) and may favor another

form of stress relaxation over whisker growth, like massive bulk deformation.10

The whiskers which formed on the surface of the three Bi-containing alloys all differed

from those seen on SAC305. While all four alloys showed a small amount of hillock

growth, particularly in locations 1, 2 and 4, the Bi containing alloys also showed a

significant amount of small whisker nucleation sites. Further study is needed to quantify

this behavior, however it is believed that stress relief through many, short whiskers

presents less of an overall reliability concern. If sufficient stress is relieved through many

short whiskers, it is believed that the likelihood of one whisker continuing to grow to a

catastrophic length may be reduced.

Another observation of this study worthy of further consideration, is the local presence of

Bi at the site of whisker growth. It was found that the final composition of the solder,

which is a combination of the solder paste, the tin plating on the component and any

diffused Cu from the lead and board, is not uniform along the entire length of the lead. In

areas which have the greatest risk of whisker growth, those along the upper lead where

the solder is thin, also have the lowest overall composition of Bi. In solder joints formed

with Senju M42, which has an initial concentration of 3wt% Bi, the Bi concentration at

areas of whisker growth was experimentally found to be between 1.3 and 1.8wt%. Solder

joints formed from Sunflower and Sunrise, which have an initial concentration of 7wt%

Bi had between 2.9 and 4.1 wt% Bi in areas of whisker growth.

Finally, in Sunrise and Sunflower, each of which have 7wt% Bi, the Bi was found to have

precipitated from the Sn, likely from the interdendritic eutectic region. This process was

shown, in section 6.1.1, to have started during solidification. Over the course of the

thermal cycling, the Bi particles have been observed along the grain boundaries.

Additionally, very fine precipitates appear to have formed during the course of thermal

cycling. Therefore it is not known what the overall impact of Bi content in the solders is

on the overall propensity for whisker growth.

154

5 References

1 JP002: Current Tin Whisker Theory and Mitigation Practices Guideline, March 2006.

2 P.T. Vianco “Dynamic Recystallization (DRX) as the Mechanism for Sn Whisker

Development. Part I: A Model” ” in Journal of Electronic Materials, Vol. 38, No.

9, 2009.

3 L. Panashchenko “Evaluation of Environmental Tests For Tin Whisker Assessment”

M.Sc. Thesis, Dept. Mech. Eng., University of Maryland, College Park,

Maryland. 2009.

4 JESD22-A121A: Test Method for Measuring Whisker Growth on Tin and Tin Alloy

Surface Finishes, July 2008

5 JESD201A: Environmental Acceptance Requirements for Tin Whisker Susceptibility of

Tin and Tin Alloy Surface Finishes, September 2008

6 P.Snugovsky, S. Meschter, Z. Bagheri, E.Kosiba, M.Romansky, J. Kennedy “Whisker

Formation Induced by Component and Assembly Ionic Contamination” in the

Journal of Electronic Materials, Vol. 41, No. 2, 2012.

7 P.Snugovsky, Z. Bagheri, M. Romansky “Whisker Growth on SAC Solder Joints:

Microstructure Analysis” in ICSR SMTA Conference, Toronto, ON, 2008

8 IPC-A-610E-2010: Acceptability of Electronic Assemblies

9 K-N. Tu “Spontaneous Tin Whisker Growth: Mechanism and Prevention” in Solder

Joint Technology : Materials, Properties and Reliability New York, Springer,

2007, ch. 6, pp. 153-181

10 S.J. Meschter, P. Snugovsky, J. Kennedy, Z. Bagheri, and E.Kosiba “Strategic

Environmental Research and Development Program (SERDP) Tin Whisker

Testing and Modeling: Thermal Cycling Testing” presented at International

Conference on Solder Reliability, Toronto, Ontario, Canada, 2014.

11 P.T. Vianco “Dynamic Recystallization (DRX) as the Mechanism for Sn Whisker

Development. Part II: Experimental Study” in Journal of Electronic Materials,

Vol. 38, No. 9, 2009.

12 P.Snugovsky, S. Meschter, Z. Bagheri, E. Kosiba, M. Romansky, J. Kennedy “Whisker

Formation on SAC305 Assemblies” in TMS Conference, San Diego CA, 2014

13 S.J. Meschter, P. Snugovsky, J. Kennedy, Z. Bagheri, and E. Kosiba “SERDP Tin

Whisker Testing: Low Stress Conditions” presented at International Conference

on Solder Reliability, Toronto, Ontario, Canada, 2012.

155

14 D.Kwon “Packaging Architecture and Assembly Technology” in A Guide to Lead-free

Solders: Physical Metallurgy and Reliability, Silver Spring, MD: Springer, 2005,

pp. 35-36

15 A.Baated , K-S. Kim, K. Suganuma, S. Huang, B. Jurcik, S. Nozawa, B. Stone, M.

Ueshima. “Effects of Reflow Atmosphere and Flux on Tin Whisker Growth of

Sn-Ag-Cu Solder” presented at SMTAI, Chicago, Il, 2009

16 P.Snugovsky, E. Kosiba, S. Meschter, Z. Bagheri, J. Kennedy “Assembly Cleanliness

and Whisker Formation” presented at IPC APEX EXPO Conference, San Diego,

CA, 2015.

17 M.Osterrnan “Mitigation Strategies for Tin Whiskers” prepared for CALCE Working

Group, 2002.

18 J-L Jo “Tin Whisker Growth Mechanism and Mitigation for Lead-Free Electronics”

Ph.D. Thesis, Dept. of Adaptive Machine System, Osaka University, Japan. 2013.

19 N.Jadhav, M. Williams, F.Pei, G. Stafford, E. Chason. “Altering the Mechanical

Properties of Sn Films by Alloying with Bi: Mimicking the Effect of Pb to

Suppress Whiskers” ” in Journal of Electronic Materials, Vol. 42, No. 2, 2013.

20 W.J. Bottinger , C.E. Johnson, L.A. Bendersky, K.-W. Moon, M.E. Williams, G.R.

Stafford “Whisker and Hillock Formation on Sn, Sn-Cu and Sn-Pb

electrodeposits” in Acta Materialia, Vol. 53 pp. 5033-5050, 2005.

21 GEIA-HB-0005-2: Standard for Mitigating the Effects of Tin Whiskers in Aerospace

and High Performance Electronic Systems, January 2007.

22 S.Meschter, P. Snugovsky, J. Kennedy, S. McKeown and E. Kosiba “Tin Whisker

Testing and Risk Modeling Project” SMTA Journal Volume 24, Issue 3 (2011)

156

Chapter 5 Mechanical (Drop) Shock Testing

1 Introduction

The proliferation of hand held devices and complex portable electronic devices occurred

at the same time as the requirement of Pb-free solders took effect.1 This confluence of

usage and regulation has introduced some significant issues related to the drop shock

response of new Pb-free solder alloys. These devices are particularly susceptible to

accidental drops over the course of their life time. At the same time SAC305 solder,

which is currently the favorite amongst commercial product manufacturers, is

significantly stiffer than traditional SnPb solder, and therefore tends to perform poorly in

response to drop shock.

As described in Chapter Chapter 1, new alloys are being explored with the intent of

improving upon SAC305. Within a SAC alloy, the Sn phase exhibits the lowest elastic

modulus and yield strength, and is therefore the most ductile phase within the SAC

solder. The IMCs which form within the bulk solder act to increase the overall strength

and reduce the ductility of the solder. In drop shock conditions, this higher strength and

lower acoustic impedance allow for the stress to more readily transfer to the solder-

copper interface, the more brittle IMC layer.2 Brittle fracture along the IMC is the typical

failure mode in drop shock conditions of a SAC solder joint as opposed to a more ductile

type fracture of SnPb solder joints, which typically failed within the bulk solder or a

combination of bulk solder and interface fracture.3 Additionally,

an increased amount of

Ag can lead to the formation of platelet shaped IMC, as opposed to the finer particles

found with low concentrations of Ag. These platelets may act as stress concentrators,

reducing the overall performance of the solder joint in drop shock testing. It has been

shown that improved drop shock performance can be attributed to lowering of the Ag

component of an alloy, i.e. from SAC405 to SAC105. This transition to better

performance occurs around the 3% Ag level, alloys with Ag below 3% perform better

157

than those with Ag above 3%.4 Lowering the Ag content will reduces the yield strength

and consequently increases the bulk solder joints ability to dissipate high plastic energy.1

Finally the higher process temperature required for proper formation of SAC solder joints

presents two additional problems, which impact the solder joints resistance to drop shock:

increased exposure to higher temperatures may lead to a thickening of the brittle IMC

interface layer abdrequiring higher Tg board materials to withstand the higher

temperature. These new board materials are more prone to a failure mode which has not

been typically seen with SnPb alloys – pad cratering. Any new alternative alloy will

therefore need to address both failure modes.

Bi as an alloying element has typically been used to reduce the melting temperature of

the main alloy. Bi may also suppress the formation of Ag3Sn platelets and reduce

intermetallic compound (IMC) growth. Furthermore, Bi acts to refine the grain structure

of the bulk solder when added to a SAC solder joint.2

This work considers a number of low (or no) Ag Pb-free alloys with differing degrees of

Bi. All of these alloys have melting temperatures between 10 and 18C lower than that

of SAC305 allowing for the use of a lower Tg board material. It is believed that the

improved properties of the new, low melt alloys, along with the lower Tg board material

will have a combined effect of improving the mechanical strength of the overall solder

joint.

In this screening experiment the drop shock testing was performed on boards with an

OSP finish. It should be noted that the board finish, and subsequent interconnection IMC,

may greatly influence the performance of an alloy. Future testing should therefore be

performed on other surface finishes.


The work presented in this section forms a screening experiment in which the selected

alloys were screened for drop shock response; the results are intended to determine

which, if any, of these alloys would be good candidates for further testing.

158

2.1 Materials

Two alloys with low Ag (Senju M42 and Sunrise), and one alloy with no Ag (Sunflower),

and varying amounts of Bi, were tested against SAC305. The alloy selection method was

described in Chapter Chapter 1. The board material was also a factor in these screening

experiments. Two board materials were tested; a normal Tg material– one which is

typically used for SnPb applications around 150C, and a high Tg material– one which is

now required for most Pb-free applications at around 170C.

Table 41: Build matrix for drop shock testing

Alloy Composition Board Material Assembly Temperature


High Tg

(170C)

240C




Sunrise Sn 1%Ag 0.7%Cu 7%Bi Normal Tg

(150C)

222C


2.2 Test Vehicle

Celestica’s RIA3 test vehicle, shown in Figure 113, was used. It was originally designed

to simulate a typical, medium complexity assembly. It is an 8”x10” PWB made up of 12

copper layers for a total thickness of 0.093” with an Organic Solderability Preservative

(OSP) finish. This board is one which is often used to test new, lead-free solders and

other material parameters. LQFP176, PBGA256, CBGA64 and MLF20s were populated,

two of each on each board. The BGA components all had SAC305 ball alloy and the

reflow was performed in an air environment. While this test vehicle was not originally

designed for evaluating consumer electronics, it served the needs of the screening

experiment in that it provided for monitoring a variety of component types.

159

Figure 113: Test vehicle with monitored components

Pad cratering has been identified as the main failure mode for Pb-free solders in

mechanical testing. This is due to the changes in laminate materials required to survive a

higher reflow temperature. PWB manufacturers have had to change the epoxy resins as

well as increase the amount of ceramic particle filler material in order to reduce the

coefficient of thermal expansion (CTE).5 To mitigate this failure mode in the drop shock

portion of testing, Solder Mask Defined (SMD) BGA pads were used. Figure 114

illustrates how the SMD boards introduce a sharp corner to the solder joints. It was

intended that this corner act as a stress concentrator within the solder joint and drive any

failure into the solder as opposed to within the laminate material, thereby allowing for a

comparison of the different solder alloys. SMD pads are also used in many commercial

applications such as cell phones and are therefore representative of commercial products

where drop strength would be a requirement.

160

Figure 114: Solder mask defined vs. non-solder mask defined6

2.3 Assembly

The assembly process utilized surface mount technology (SMT) parts, as described

above, placed and then reflowed with target reflow profiles outlined in Table 41. Primary

SMT was performed with No Clean flux in a ten-zone oven. 42 RIA3 boards were built

with an OSP finish and solder mask defined (SMD) copper pads. The RIA3 card was cut

in half in order to better represent the smaller board sizes used in commercial applications

(i.e. smart phones). One half of the board (RIA3-1) had five monitored components,

while the other half (RIA3-2) had only one monitored component as per Table 42 and

Figure 113.

2.4 Test Strategy

Five boards of each combination were exposed to drop shock testing, one was used for

Time 0 analysis, and one was held back as a “do nothing” board for possible verification

purposes which may arise in the future.

Table 42: Monitored components for drop testing

Board Component Reference Designator

RIA3-1 BGA-256 U204

RIA3-1 BGA-256 U205

RIA3-1 QFP-176 U1

RIA3-2 QFP-176 U2

RIA3-1 CSP-64 U208

RIA3-1 CSP-64 U209

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In order to test the mechanical strength of the solder joints, board level drop testing,

based on JESD22-B110A7, was performed. This test method aims to evaluate a

subassembly’s ability to “withstand moderately severe shocks as a result of suddenly

applied forces or abrupt change in motion”. Within a subassembly there are four basic

types of failure which typically resulting from drop shock:

1) Fracture, or permanent deformation, caused by high applied stress

2) Chatter created by high acceleration levels, e.g. causing bolts to loosen

3) Impact between adjacent objects, caused by high displacement

4) Momentary electrical failure associated with a shock pulse8

In this test, condition 1 is the most likely and desired failure mode. Conditions 2 and 3

are mitigated by the test set up in which all resistance cables, strain gages and

accelerometers are secured using room temperature vulcanization (RTV) silicone, as seen

in Figure 115, and loaded within a fixture allowing for sufficient spacing to avoid impact

with other objects as seen in Figure 116. Condition 4 presents a problem as a momentary

electrical failure, which recovers after a shock pulse, and may be the result of an

electrical response. This is mitigated by using an event detector to monitor failures,

which require an increase of 300 for at least 200ns. An electrical pulse less than this

would not be recorded as a failure.

162

Figure 115: Example of accelerometer secured with RTV silicone

Figure 116: Test set up

As per JESD22-B110, the subassembly was supported in a manner which is

representative of field conditions, in this case allowing the PWB to flex during the drop

shock. The parameters are a peak acceleration of 1500G with a 0.5ms duration defined in

Table Accelerometer

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JESD22-B110 Service Condition B in a half-sine waveform. Figure 117 and Figure 118

show the target and the achieved half-sine waveform respectively. In Figure 118 the red

line represents the acceleration experienced by the drop table, monitored by an

accelerometer attached directly to the table. A limited amount of rebound was measured

immediately after the drop. This may be attributed to the movement of the accelerometer

relative to the board (or table) to which it is affixed, which should be minimized by use of

RTV silicone or a noise in the collected signal. It may also be the result of poor

dampening. Compared to the initial drop shock signal, this rebound is relatively small

and therefore not considered a significant contributing factor. The green line in Figure

118 represents the acceleration experienced by the PWB itself and is monitored via an

accelerometer attached directly to the board, generally directly opposite the monitored

component shown in Figure 113. As the card is secured to the fixture by posts at the four

corners of the board, it was allowed to freely flex back and forth after the initial shock

impulse. Per JESD22-B110, this back and forth flexure is representative of a subassembly

life condition and therefore a desired part of the test set up. An Analysis Tech STD-256

event detector was used to monitor the resistance threshold of the components under test

during the mechanical shock. A failure was recorded when the channel resistance

increased by 300 or more for at least 200ns. The testing was stopped as soon as an

electrical failure was detected (first failure).

Figure 117: Target pulse shock defined by JESD22-B110 service condition B

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Figure 118: Sample of pulse shock achieved during test

While the method outlined for this test can be used to calculate the change in velocity and

the displacement of the monitored items, in this project it was used to compare the

response of the various alloys. Specifically, the number of drops required to produce an

electrical failure was measured, while all other parameters were monitored to ensure

consistency of test conditions.

3 Reliability Results

Drop shock testing as described above was performed on five boards of each combination

described in Table 41. Each of the two halves was tested separately. It was initially

intended that all components would be monitored to failure allowing for 10 failures to be

recorded for each component type. This would provide a small sample size, but one

sufficient to produce basic Weibull plots. A number of events occurred during the testing

which reduced the number of components tested. These events are outlined below:

Based on similar tests performed in the past, it was deemed that additional weights

were needed to increase the strain and help induce failure in the solder joints. Two

220g weights were added to the center of each board.9 After testing five cards, it

was determined the weight was not needed to induce failure. The resulting data

165

associated with RIA3-1 portion of five boards was not used in the analysis. Testing

then continued with no additional weight added to the remaining boards.

The process of attaching the weights described above involved drilling into the

laminate material to securely attach the additional weight. The resulting hole was

located very close to U1 on RIA3-1. It was therefore decided not to use any data

obtained from the monitoring of U1 as it could not be determined whether the

failure was due to the drilling process or the drop shock, or a combination of the

two. No drilling was performed on RIA3-2 and therefore all data from U2 was used

in this evaluation.

The CSP components, as well as the BGA, did not fail. These components were

located towards the edge and corners of the test vehicle and would therefore not

experience the same degree of strain as those located in the centre of the test

vehicle.

Therefore, data from a total of 25 out of 30 BGAs and from all 30 QFPs was available for

evaluation. The results are given in Table 43 and in Figure 119 and Figure 120.

166

Table 43: Drop test results

Board Material

(Tg) Alloy Board Number

Drops to Failure

BGA (U205) QFP (U2)

High Tg

(170C)

SAC305

22 --* 66

24 17 65

26 25 63

27 16 52

28 8 72

Senju M42

1 16 56

2 19 42

3 10 60

4 9 37

6 11 51

Sunrise

15 9 38

16 --* 45

19 15 62

20 8 50

21 --* 56

Sunflower

8 7 55

10 --* 57

11 14 46

13 20 65

14 10 46

Normal Tg

(150C)

Sunrise

36 23 59

38 8 82

40 26 56

41 34 57

42 8 76

Sunflower

29 20 73

30 13 63

31 11 65

34 46 79

35 --* 84

* Samples removed for reasons provided on page 164

167

Figure 119: Individual value plot of drops to fail, BGA (SAC305 + alloy) (U205)

Figure 120: Individual value plot of drops to fail, QFP (U2)

168

The results of testing, albeit with a small sample size, indicate the board material was a

more significant factor in drop shock performance rather than the alloy. Normal Tg

boards outperformed High Tg boards in all cases.

The data also indicates that SAC305 outperformed the other alloys when comparing the

High Tg board materials. This is particularly significant in the QFPs, likely due to the

fact that the QFP solder joints are made up of 100% test alloy, the BGA solder joints are

made up of 13% test alloy and 87% SAC305 from the component solder ball. There is

therefore less of a difference in the overall composition of the BGA solder joints.

Among the three test alloys, it appears that Sunflower performed the best; however there

is not enough data for a conclusive determination. A larger sample size and perhaps a

more appropriate test vehicle would be able to provide a greater distinction between the

alloys.

It is important to note that the limited reliability data provides only a cursory view,

providing information only on which solder joints failed during the drop testing. For a

better understanding of what occurred, failure analysis of the solder joints was required to

form a clearer picture of the drop shock response. Both reliability and failure analysis

need to be assessed together for a complete understanding.

4 Failure Analysis and Microstructural Evaluation

Failure analysis of the drop shock test boards was performed by one of two methods:

Dye and Pry (D&P) Testing: This method provides an overall image of the extent of

the failures across an entire component, as well as allowing for the determination of

the failure mode

Cross Sectional Analysis: This method provides a clearer observation of the failure

mode in the solder joint; however, it limits the number of solder joints to be

examined.

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4.1 Dye and Pry Procedure

Dye and Pry testing was performed per Celestica’s procedure DOC0018652 “Dye and

Pry Testing”10

. The boards were immersed in red dye and subjected to a vacuum in order

to force the dye into any pre-existing cracks, which occurred during drop testing. The

dye was cured prior to component removal. The fracture surfaces were then inspected for

the presence of dye – a percentage of dye penetration for each solder joint was

determined, as was the mode of failure (refer to Figure 121).

Figure 121: Failure modes of solder joint as defined by IPC/JEDEC-970211

4.2 Failure Isolation Procedure

The location of failures was isolated by probing the daisy chain pattern using vias on the

bottom side of the PWB for BGAs, or between leads for QFPs. Cross sections were then

prepared at locations where electrical failures were found to have occurred.

4.3 Evaluation of High Tg Board after Drop Testing

Evaluation performed using D&P and cross sectioning showed that BGAs on high Tg

boards failed exclusively by pad cratering. Figure 122 illustrates a pad crater from a cross

sectional perspective. The solder joint remains completely intact; no cracking was seen

along the IMC or within the bulk solder is observed, furthermore there was no separation

between the IMC and the copper pad. All stress was dissipated through the laminate

material, which failed between the epoxy and the layers of glass weave within the board

material.

170

Figure 122: BGA failure by pad cratering, Sunflower on High Tg boards

Dye and pry results allow for an overall view of the failures within a component, both the

number of failed solder joints and the extent to which the fracture penetrated through the

solder joint. It is important to note that pad cratering may not result in an electrical

failure. Partial fractures, as indicated by the various shades of yellow/orange in the D&P

mapping, allow for some electrical contact to be retained. Even complete fractures,

indicated by red, may still have intact copper traces associated with them. These types of

failures are particularly insidious because they are not readily identifiable but do present

a real reliability threat as the soft copper trace continues to be stressed. Figure 123 shows

the D&P mapping of one component with accompanying images of representative

failures in specific solder joints. The number within each cell of the map corresponds to

a failure mode described in Figure 121. Table 44 provides a summary of the failures, as

seen by D&P of all the alloys tested on High Tg board material. % Fracture corresponds

to the total area of the fracture surface, which was penetrated by dye. This indicates the

degree to which the fracture penetrated through the solder joint prior to the pry process.

The next four columns show the number of solder joints, as a percentage of the solder

joints on the component, which showed some degree of dye penetration. From this we see

that Senju M42 showed the least amount of damage. All of the components selected for

D&P failed between 8 and 10 drops and all showed complete dye penetration within

many of the solder joints on the outer edge of the component closest to the center of the

171

board. All solder joints which did not show any dye penetration resisted fracture up to

the point of first failure.

Figure 123: D&P mapping of Senju M42 on High Tg board, with images of

board side and b) component side of the fracture surface

Table 44: Failures of BGA on High Tg boards

SAC305 Senju M42 Sunrise Sunflower

75-100% 7% 4% 7% 6%

50-75% 5% 1% 3% 4%

25-50% 7% 5% 9% 7%

5-25% 5% 1% 4% 5%

Total: 24% 11% 23% 21%

High Tg Board% Fractured

The QFPs on high Tg boards did not exhibit any pad cratering. This was likely due to the

structure of the part, as illustrated in Figure 124, which is much more compliant than a

BGA. With copper leads on all four sides, the component itself is able to efficiently

dissipate stress.12

Electrical failures were mainly associated with damaged leads, as seen

in Figure 125a, which after a number of drops failed at the “knee” where large repetitive

172

strain would have occurred. Although no complete mechanical failure through the bulk

solder was observed on QFPs, the beginning of IMC failure was observed in a number of

the solder joints. The crack would begin on the outside of the solder fillet and move to

the copper lead where it would continue to propagate along the brittle IMC layer as seen

in Figure 125b.

Figure 124: Schematic of Quad-Flat-Package (QFP)13

Figure 125: QFP failures in Sunflower on High Tg boards

a) fractured lead b) solder fillet fracture

173

4.4 Evaluation of Normal Tg Board after Drop Testing

Sunrise and Sunflower alloys were tested on Normal Tg boards, as seen in Figure 119

and Figure 120. Both alloys survived, on average, more drops to first failure than on High

Tg boards. Failure analysis also revealed differences in the failure mode and the degree

of damage occurring within the component and within each individual solder joint.

Figure 126 shows the D&P mapping of Sunrise on a Normal Tg board. While a small

amount of pad cratering was visible on one edge of the component, the dominant failure

mode was between the IMC interfaces with the bulk solder material towards the board

side of the component. Figure 127 shows this same failure mode, between the IMC and

bulk solder, as seen in a cross section of a BGA solder joint formed between SAC305

(BGA ball) and Sunflower (paste). Finally, Table 45 summarizes the degree of failure

which occurred in BGAs using these two alloys. While the amount of failure appeared

relatively similar between the two alloys, it is significantly less damage than was

observed on the High Tg boards using the same alloys.

Figure 126: D&P mapping of Sunrise on Normal Tg board, with images of

board side and b) component side of the fracture surface

174

Table 45: Failures of BGA on Normal Tg boards

Sunrise Sunflower

75-100% 1% 1%

50-75% 2% 1%

25-50% 1% 0%

5-25% 4% 4%

Total: 8% 6%

Normal Tg Board% Fractured

Figure 127: BGA failure in Sunflower on Normal Tg boards

The QFPs, as in the case of High Tg boards, failed via lead fracture as seen in Figure

128a. No pad cratering was observed. Fracture, initiating in the solder fillet and

propagating along the IMC was observed, however no complete fracture of this type was

observed.

175

Figure 128: QFP failures in Sunflower on Normal Tg boards

a) fractured lead b) solder fillet fracture

The solder alloy may be a factor in failures caused by fractured leads; the cracks appear

to originate in the solder material and then propagate through the lead at a point of high

mechanical strain.14

This is illustrated in Figure 129.

Figure 129: QFP failure in Sunrise on Normal Tg board

5 Summary of Findings and Conclusion

The drop testing carried out comprised a screening experiment. While the sample size

may not be statistically significant, it does show some promising results. Many of the

results relate to the board material rather than to the solder material. This is significant as

176

lower Tg board materials are a viable option for lower melting temperature solders,

SAC305 often requires higher Tg board material.

Normal Tg boards appear to outperform High Tg boards independent of solder alloy.

High Tg boards failed exclusively by pad cratering, which is a latent defect and

therefore not readily detectable. Pad craters present a reliability concern because

they will propagate through the laminate and eventually sever the copper trace. Pad

craters will not be detectable until the copper trace is broken leading to electrical

failure. Even if the laminate defect could be detected at manufacturing, it is not

readily repairable.

Normal Tg boards survived longer. While the solder joints built on Normal Tg boards

exhibited mixed failure modes (sometimes failing by pad cratering and other times

by solder fracture at the IMC) there were also significantly less failed solder joints,

and they primarily failed at the IMC interface.

Additionally, there were some differences among the solder alloys used in drop shock

testing. On the High Tg board material, there is not enough data to properly distinguish

between the performances of each alloys in terms of drops to first fail, however Senju

M42 does appear to have sustained the least amount of damage. Furthermore, when

comparing Sunflower to Sunrise on both High and Normal Tg board materials, it appears

that Sunflower slightly outperforms Sunrise.

The following are recommendations for further testing based on this screening

experiment:

The ideal test vehicle would have one component at the centre of the board, which

could be monitored to failure.

Each solder/board material combination should include 32 repeats, of which at least

63% are tested to failure in order to provide sufficient data for reliability plots

Using both SAC305 and SnPb as a baseline for comparison would provide a better

understanding of how low-melt, Bi-containing alloys perform against current

industry standards.

177

6 References

1 D.A. Shnawah, M.Sabri, I.A. Badruddin “A review on thermal cycling and drop impact

reliability of SAC solder joint in portable electronic products”, Microelectronics

Reliability, 52 (2012) pp. 90-99

2 P. Ranjit and T. Lawlor., “Effects of Silver in common lead-free alloys”

3 T. A. Woodrow and J. Bath “Lead-Free Reliability in Aerospace/Military

Environments” in Lead-Free Solder Process Development, New York, Wiley,

2011

4 G. Henshall, R. Healey, R. Pandher, K. Sweatman, K. Howell, R. Coyle, T. Sack, P. Snugovsky, S.

Tisdale, F. Hua. “iNEMI Pb-Free Alloy Alternatives Project Report: State of the Industry,”

proceedings SMTAI, p. 109 (2008).

5 B.Gray “Correlation of Printed Circuit Board Properties to Pad-Crate Defects Under

Monotonic Spherical Bend”, M.A.Sc. thesis, Dept. Mech. Eng., Ryerson

University, Toronto, Ontario. 2012.

6 PCB Layout Recommendations for BGA Packages, Lattice Semiconductor Corp. Oct.

2005

7 Subassembly Mechanical Shock, JESD22-B110A, November 2004.

8 D.S. Steinberg “Designing Electronics for Shock Environment”, in Vibration Analysis

for Electronic Equipment, 3rd

ed. New York, Wiley, 2000. ch. 11, pp.248-298

9 W.Liu, N-C, Lee, S. Bagheri, p. Snugovsky, R.Brush, J.Bragg, B.Harper. “Drop Test

Performance of BGA Assembly using SAC105Ti Solder Spheres” in IPC APEX

EXPO Proceedings, 2011

10 Dye and Pry Testing, Celestica DOC0018652, 2014.

11 Monotonic Bend Characterization of Board-Level Interconnects, IPC/JEDEC-9702,

June 2004.

12 P. Snugovsky, J. Bragg, E. Kosiba, M. Thomson, B. Lee, R. Brush, S. Subramaniam,

M. Romansky, A. Ganster, W. Russell, J. P. Tucker, C.A. Handwerker, D.D.

Fritz. “Drop Test Performance of A Medium Complexity Lead-Free Board After

Assembly and Rework” in the IPC APEX EXPO Conference Proceedings, 2010.

13 http://www.practicalcomponents.com/Dummy-Components/




Vegas, NV, 2014.

178

Chapter 6 Summary of Findings and Conclusions

This work explored the reliability performance and microstructural differences of three

low melting temperature, low-Ag, Bi-containing alloys compared with SAC305. From

the current work, the following conclusions can be made:

The three low melting temperature, low-Ag, Bi-containing alloys under investigation

all produced good solder joints with QFP type components. BGA components with

SAC305 solder balls also formed good solder joints when built with these pastes;

all examined solder joints had an acceptable degree mixing. Differences in

microstructure have been observed related to the addition of Bi, decrease in Ag-

content, and varying amounts of Cu dissolution resulting from different process

temperatures.

Solder joints with final Bi concentrations of less the ~3wt% showed no detectable Bi

precipitates. This includes QFP solder joints formed with Senju M42 and all BGA

solder joints formed with a combination of paste and SAC305 solder balls. In these

cases, all the Bi was present within the Sn phase upon solidification.

Bi was observed to precipitate out of solid solution from “as assembled” solder joints

with Bi compositions of greater than 7wt%. This is the case with QFP solder joints

formed using Sunrise and Sunflower solder pastes. Bi also remains present in the

Sn phase, observed through EDX analysis. It is therefore likely that the Bi

precipitates are formed from the binary, ternary or double binary eutectic regions

within the interdendritic spaces of the Sn phase where the concentration of Bi in

Sn is highest. Both fine and medium sized particles were observed. In

Sunflower, the concentration of Bi within the Sn dendrites and within the Sn

portion of the binary eutectic (Sn + Cu6Sn5) within the interdendritic spaces

appeared to be approximately equal.

Bi concentrations greater then 7wt% had a beneficial impact on the changes to bulk

microstructure, which occurred during accelerated thermal cycling. The Bi

precipitated from the bulk solder evenly in very small particles. This favorable

179

influence did not appear when the Bi concentration was 3wt%. There was some

degree of Bi segregate along the grain boundary observed in the Sunflower alloy;

however there did not appear to be full segregation of the phases.

After exposure to thermal shock, Bi particles were observed along the grain

boundaries in solder joints formed with alloys containing 7wt% Bi. Additionally,

very fine precipitates appear to have formed during the course of thermal shock.

Cu6Sn5 () interfacial IMC layers that form between the studied solders and the Cu

pad of the board or the lead material of the QFP solder joint during solidification

were not statistically distinguishable in terms of thickness. No other species was

detected at the solder/Cu interface.

In the BGA solder joint, the interfacial IMC layer that formed between the combined

solder paste/SAC305 solder ball and the Cu pad of the board was found to be

(Cu,Ni)6Sn5. The Ni is contributed to this system from the component side IMC and

diffuses through the bulk solder prior to solidification. Using a 93% CI, SAC305

appeared to have a thicker IMC layer upon solidification. This can be attributed to

the higher process temperature. The interfacial IMC layer that formed between the

Cu/Ni/(paste/SAC305 solder ball) was found to be (Cu,Ni)6Sn5 and Ni23Cu33Sn44.

Due to the irregular shape of this IMC layer, it was difficult to compare the mean

thicknesses of the IMC layer that formed when using different solder paste alloys.

During thermal cycling, the total thickness of the interfacial IMC continued to

increase, and a () Cu3Sn layer formed. While the overall thickness increased

similarly for all four alloys, the Cu3Sn layer of the Sunflower alloy was much

thicker than the other alloys. This alloy has no Ag, which indicates that Ag plays a

role in suppressing Cu3Sn growth. Bi concentration did not appear to significantly

change the growth rates of either intermetallic compound. The increased growth of

the Cu3Sn phase in the Sunflower alloy did not correspond to any decrease in

reliability during ATC.

180

SAC305, which was processed at a higher temperature, initially had a slightly thicker

interfacial IMC layer. During thermal cycling, the IMC layers formed in solder

joints using the various alloys increased to similar overall thicknesses.

The three lower process temperature solders performed as well as, or better than

SAC305 during accelerate thermal cycling. During two different thermal cycling

regimes (0 to 100C and -55 to 125C) the only failures which could be attributed

to solder joint fatigue failures occurred in SAC305 during 0 to 100C cycling and

one sample of Sunrise, which failed towards the end of the test. All other electrical

failures were attributed to failures within the board material. High Tg board material

appeared to outperform Normal Tg board material during ATC.

High Temperature, High Humidity conditions (85C, 85%RH) produced an

insufficient number of whiskers and no whiskers which would fail the requirements

of JEDEC Standard No. 201A for class 2 components. This test was run for 1000

hours, while current required length of test is 3000 hours. It would therefore be

worthwhile to continue testing.

1610 cycles of Thermal Shock testing is sufficient to meet the requirements for

whisker testing. In this case, all four alloys passed the requirements set out in

JEDEC Standard No. 201A for class 2 components.

The whiskers which formed on the surface of the three Bi containing alloys during

thermal shock all differed from those seen on SAC305. All alloys showed a small

amount of hillock growth, the Bi containing alloys also showed a significant

amount of small whisker nucleation sites. This may prove beneficial as many short

whiskers may effectively reduce the stress in the system and decrease the likelihood

of one whisker continuing to grow to a catastrophic length.

The final composition of the solder, formed between the solder paste and the

dissolving Sn finish, is not uniform along the entire length of the lead on a QFP

component. In areas which have the greatest risk of whisker growth, those along

the upper lead where the solder is thin, also have the lowest overall concentration of

Bi.

181

Normal Tg boards appear to outperform High Tg boards in drop testing, independent

of solder alloy. This is significant as lower Tg board materials are a viable option

for lower melting temperature solders where as SAC305 often requires higher Tg

board material. The manner in which boards fail varies according to board Tg: High

Tg boards failed exclusively by pad cratering, Normal Tg boards, which survived

longer, than High Tg boards, exhibited mixed failure mode of pad cratering and/or

brittle fracture at the IMC.

While there was not enough data to distinguish the different solder paste alloys based

on number of drops to failure, there was some interesting information gleaned

during failure analysis regarding the degree or number of solder joints which failed

per component. When analysing High Tg boards, Senju M42 appears to have

sustained the least amount of damage. The other three alloys were not

distinguishable. Furthermore, when comparing Sunflower to Sunrise on both High

and Normal Tg board materials, it appears that Sunflower slightly outperforms

Sunrise. Significantly fewer solder joints per component failed when using Normal

Tg boards.

Based on comparisons to the mainstream consumer alloy –SAC305, which has a

minimum reflow temperature of 223C and a Ag concentration of 3wt%, the new, low

melting temperature, low-Ag, Bi-containing alloys may be recommended as viable

replacements.

182

Chapter 7 Future Work

The work in this thesis, as well as that in the proposed future work, is a continuation of

the research collaboration between Celestica Inc and U of T into the development of a

reliable solder alloys, that will conform with the requirements of their respective end use

environments while meeting the environmental lead (Pb)-free constraints imposed on the

commercial supply base.

1 Celestica/Indium Sponsored Whisker Resistant Solder Paste

It has been shown that component and assembly cleanliness affect the propensity of a

solder joint to grow whiskers, particularly in high humidity environments. It was found

that contamination impacted the degree to which corrosion formed at the surface, and

consequently the extent of internal compressive stresses applied at the surface of the

solder, which is a required element in whisker nucleation and growth.1 In the production

process, a major source of contamination is the flux residue, which may remain on the

product after the build. The role of flux is to provide an oxide-free surface for the solder

to bond with a base metal, usually copper. It is intended that the flux should be fully

consumed, or burned off, by the end of the reflow process. In practice however, it is

possible that some residue remains after production. This flux residue now behaves as a

contaminate. Various flux types have been formulated to reduce the corrosive impact of

the flux residue depending on the end use of the electronic product.

In addition to RoHS legislation, the electronics industry is undertaking further

environmental initiatives, including the reduction of halogens from electronic products.2,3

This initiative restricts the use of chlorides and bromides throughout the electronic supply

chain including the solder paste flux formulations. Halides have been used in fluxes in

order to rapidly reduce the metal oxide and allow for the formation of a robust solder

joint. Any remaining halide may, however, behave as a corroding agent if not fully

183

utilized in the reflow process. In a “halogen-free” solder paste, it is still unclear whether

the alternative agents will be more, or less corrosive.

In developing solder pastes for low melt applications with three to four elements (Sn, Cu,

Bi and sometimes Ag), Indium and Celestica are working to optimize the solder flux as

well as the alloy. It is therefore important to identify the corrosion mechanism(s) which

may affect the various elements and compounds within a particular solder. For example,

it is hoped that this work will determine whether the dominant form of corrosion is

intergranular, galvanic or some other form.

23 flux/paste combinations will be tested according to the matrix outlined in Table 46

with the goal of better understanding the interaction of the flux/paste combination.

Table 46: Solder paste test matrix

Flux

Alloy

SAC305

Sn 3%Ag 0.5%Cu Violet

Sn 2.25%Ag 0.5%Cu 6%Bi Sunflower

Sn 0.7%Cu 7%Bi

Base BaA1 BaA2 BaA3

Resistively 1 R1A1 R1A2 R1A3

Resistively 2 R2A1

MPU 1 M1A1 M1A2 M1A3

MPU 2 M2A1

Corrosion 1 C1A1 C1A2 C1A3

Corrosion 2 C2A1

pH (1) P1A1

pH (2) P2A1 P2A2 P2A3

pH (3) P3A1 P3A2 P3A3

pH (4) P4A1

184

2 ReMap M1: Lower Temperature Soldering Alloys with Improved Mechanical and Thermal Fatigue Reliability

The ReMap M1 project has two main goals: to develop a test protocol for testing alloys in

a combined thermal mechanical reliability test, and to evaluate Bi containing alloys

according to this protocol.

Currently, reliability tests exist for thermally induced stress (i.e. thermal cycling) and for

mechanical stress (e.g. vibration testing, drop testing) separately. There is currently no

existing industry standard test protocol for examining the interaction of these stress types.

Electronic assemblies, particularly those intended for high reliability applications, are

often exposed to a various stresses and often at the same time. It is believed that the

material behavior resulting from thermal stress will impact the material response to

mechanical stress. The mechanical behavior at various temperatures is therefore thought

to be significantly different. Also, the behavior of the material as it cycles through

different thermal regimes may also affect the response to mechanical stresses.

It is believed that the inclusion of Bi in these alloys will result in material strengthening

by one or more of the following mechanisms: solid-solution strengthening or

precipitation hardening. The mechanism will likely dependent on the amount of Bi

present in the alloy.4

The aim of this project is to correlate reliability data with the metallurgical findings in

order to better understand the behavior of various alloys under combined stress

conditions.

3 ReMap M2: Alloys, Board and Component Surface Finish Interactions with Reduced Propensity for Whisker Growth

The current work presented a preliminary assessment of the whisker mitigation properties

of a number of Bi containing alloys. This will be further developed in the ReMap M2

185

project, in which a number of Bi containing alloys will be tested (Table 47), and assessed

according to statistically accepted guidelines, in a variety of conditions, for whisker

growth characteristics. These findings will also be compared to the microstructural

properties of the alloys, as they have evolved in the listed conditions.

In previous work, it has been shown that internal stresses are created from a variety of

sources. High Temperature High Humidity (HTHH) and Ambient Temperature High

Humidity (ATHH) where selected to evaluate how internal, compressive stresses are

created at the surface of a solder joint by oxidation. This has been shown to affect Cu

leaded solder joints more than Alloy 42. ATHH with contamination is an extension of

the evaluation by oxidation, which is further enhanced by an ion-contaminated

environment. Thermal cycling, at two different levels, will also be evaluated in order to

study the affects of internal compressive stresses created from a mismatch of CTE

amongst the various materials within a solder joint. It has been shown that Alloy 42 is

affected more than Cu based lead frame alloys due to the larger difference in

CTE.5,6,7,8,9,10

It is believed that Bi mitigates the nucleation and growth of whiskers in a variety of ways.

Even a small amount of Bi has been shown to lead to a more equiaxed grain structure.

This is thought to elevate some of the internal compressive stresses on specific, columnar

grain boundaries, which have been observed in plating material. Bi also acts as a grain

refiner, by providing more grain boundaries, there is the possibility to annihilate

dislocations and therefore absorb stress internally. Finally, with the lower process

temperatures of Bi containing alloys, the scallop shaped interfacial IMC layer is likely to

be suppressed. Due to its irregular shape, this IMC layer is a major contributor to the

internal compressive stress of a thin film or a thin layer of solder.11,12

186

Table 47: ReMap M2 test matrix

Surface

Finish Alloy

Test Condition

HTHH

(85C/85% RH)

ATHH

(25C/85% RH)

ATHH

(25C/85% RH)

+

contamination

Thermal

Cycling

(-55C to

80C)

Thermal

Cycling

(-20C to

100C)

ENIG

SAC305 3x 3x 3x 3x 3x

Senju M42 3x 3x 3x 3x 3x

Sunflower 3x 3x 3x 3x 3x

Violet 3x 3x 3x 3x 3x

OSP

SAC305 3x 3x 3x 3x 3x

Senju M42 3x 3x 3x 3x 3x

Sunflower 3x 3x 3x 3x 3x

Violet 3x 3x 3x 3x 3x

4 ReMap M3: Aging Effect of New Lead-Free Materials on Reliability

The current work has focused on understanding the properties of Bi containing alloys in

the short term, with respect to time from manufacture. In reality, all of the testing took

place over 4 years and started approximately one year after the initial build, during which

time all samples were stored at ambient conditions. The effects of aging in this study are

therefore not well understood.

It has been shown in a number of studies that both SnPb alloys and SAC based alloys

exhibit significant deterioration of mechanical properties upon aging. This has been

found to occur at both room temperature and elevated storage conditions, with SAC

based solders exhibiting a deterioration of mechanical properties in the order of 25 times

greater than that experienced by SnPb.13

It is believed that the microstructural coarsening

is the main mechanism responsible for the reduction in mechanical properties; coarsening

187

occurs on both SnPb and SAC based alloys (Figure 50,Figure 51)

14. In previous work, it

appears that Bi containing alloys may actually improve the mechanical properties by

reducing the amount of microstructural coarsening and through either solution hardening

(in the case of low concentrations of Bi) or precipitation hardening (in the case of higher

Bi content).15

Figure 130 illustrates the microstructural change that occurred in Orchid

(Sn2%Ag7%Bi) after 3000 cycles of harsh environment ATC (-55C to 125C).

Significant growth of the Sn dendrites was not observed, and while the other IMC

particles did appear to coalesce to some degree, the main observed change is that the Bi

precipitated out of the Sn, likely from the interdendritic eutectic first, and remained as

small, evenly spaced precipitates.

Figure 130: Orchid at a) Time 0 and b) after 3000 cycles ATC

Lower melting point alloys, specifically those containing Bi, may also improve the

mechanical properties of a solder joint by suppressing the growth of the interfacial IMC

layer. The increased reflow temperatures associated with SAC alloys, as well as the

continuous growth of the interfacial IMC over time, has been identified as a major

contributing factor to the reduced reliability of these alloys.16

While these initial findings are promising, more study is required. Potential pitfalls of Bi

precipitation also exist. If the Bi begins to coalesce into larger particles, it will no longer

present the benefits associated with a small evenly spaced particles. Also, there is a

danger that the Bi will preferentially locate along grain boundaries or along the interfacial

IMCs, similar to the known issue with SnPb solders discussed in Section 3.

188

During the course of this study, unexpected behavior was observed. Approximately two

years after the High Temperature High Humidity exposure described in Section 2.2,

samples, which had been stored at ambient conditions, where re-examined for further

whisker growth. While a small amount of whisker nucleation was observed on the sides

of the leads, a more interesting observation was that the Bi appeared to be precipitating

out of the solder (Figure 131).

Figure 131: Sunrise solder joint after 1000 hrs HTHH followed by 2 years ambient storage

Through the ReMap M3 project, the team will investigate the evolution of three Bi

containing alloys through a number of different aging conditions (Table 48). The

microstructure and hardness of each alloy, exposed to each different test condition will be

evaluated. The samples will then be down selected for reliability study. A subset of

alloys and aging conditions will be tested in both ATC and vibration conditions, and

possibly a combined environment as described in Section 2.

189

Table 48: ReMap M3 test matrix

Surface

Finish Alloy

Aging Conditions

25 hours 300 hours

25C ~60-70C 125C 25C ~60-70C 125C

ENIG

SnPb 3x 3x 3x 3x 3x 3x

SAC305 3x 3x 3x 3x 3x 3x

Sunflower 3x 3x 3x 3x 3x 3x

Senju M42 3x 3x 3x 3x 3x 3x

Orchid 3x 3x 3x 3x 3x 3x

OSP

SnPb 3x 3x 3x 3x 3x 3x

SAC305 3x 3x 3x 3x 3x 3x

Sunflower 3x 3x 3x 3x 3x 3x

Senju M42 3x 3x 3x 3x 3x 3x

Orchid 3x 3x 3x 3x 3x 3x

190

5 References

1 P.Snugovsky, S. Meschter, Z. Bagheri, E.Kosiba, M.Romansky, J. Kennedy “Whisker

Formation Induced by Component and Assembly Ionic Contamination”, Journal

of ELECTRONIC MATERIALS, Vol. 41, No. 2, 2012. pp. 204-223.

2 Directive 2002/95/EC of the European Parliament and of the Council of 27 January

2003 on the restriction of the use of certain hazardous substances in electrical and


3 R. Flinders “Halogen-Free Electronics: An Overview” presented at IEEE PSES Orange

County Chapter Meeting Irvine, CA August 26, 2008

4 P.Snugovsky “ReMap M1 Project Proposal: Lower Temperature Soldering Alloys with

Improved Mechanical and Thermal Fatigue Reliability” April 2014.

5 P.Snugovsky, Z. Bagheri, M. Romansky “Whisker Growth on SAC Solder Joints:

Microstructure Analysis” in ICSR SMTA Conference, Toronto, ON, 2008.

6 S.J. Meschter, P. Snugovsky, J. Kennedy, Z. Bagheri, and E.Kosiba “Strategic

Environmental Research and Development Program (SERDP) Tin Whisker

Testing and Modeling: Thermal Cycling Testing” presented at International

Conference on Solder Reliability, Toronto, Ontario, Canada, 2014.

7 P.Snugovsky, S. Meschter, Z. Bagheri, E. Kosiba, M. Romansky, J. Kennedy “Whisker

Formation on SAC305 Assemblies” in TMS Conference, San Diego CA, 2014

8 S.J. Meschter, P. Snugovsky, J. Kennedy, Z. Bagheri, and E. Kosiba “SERDP Tin

Whisker Testing: Low Stress Conditions” presented at International Conference

on Solder Reliability, Toronto, Ontario, Canada, 2012.

9 P.Snugovsky, E. Kosiba, S. Meschter, Z. Bagheri, J. Kennedy “Assembly Cleanliness

and Whisker Formation” presented at IPC APEX EXPO Conference, San Diego,

CA, 2015.

10 S.Meschter, P. Snugovsky, J. Kennedy, S. McKeown and E. Kosiba “Tin Whisker

Testing and Risk Modeling Project” SMTA Journal Volume 24, Issue 3 (2011)

11 N. Jadhav, M. Williams, F. Pei, G. Stafford, E. Chason “Altering the Mechanical

Properties of Sn Films by Alloying with Bi: Mimicking the Effect of Pb to

Suppress Whiskers” ” in Journal of Electronic Materials, Vol. 42, No. 2, 2013.

12 J-L Jo “Tin Whisker Growth Mechanism and Mitigation for Lead-Free Electronics”

Ph.D. Thesis, Dept. of Adaptive Machine System, Osaka University, Japan. 2013.

13 H. Ma, J.C. Suhling, Y. Zhang , P. Lall, M.J. Bozack, “The Influence of Elevated

Temperature Aging on Reliability of Lead Free Solder Joints”, in ECTC

Conference, Sparks, NV, 2007

191




Vegas, NV, 2014.

15 P.Snugovsky. “ReMap M3 Project Proposal: Aging Effects of New Lead-Free

Materials on Overall Reliability” April 2014.

16 Z. Hai, J.Zhang, C. Shen, J.L.Evans, M.J.Bozack ”Long-Term Aging Effects on

Reliability Performance of Lead-Free Solder Joints”, presented at SMTAI,

Chicago, Il, 2013.

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