REFERENCE

NBS TECHNICAL NOTE 726

Testing and Fabrication of

Wire-Bond Electrical ConnectionsA Comprehensive Survey
NATIONAL BUREAU OF STANDARDS

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Eli.!, L_. J L'r vji.-"-!*'"^*

111372 ^^^

Testing and Fabrication of

Wire-Bond Electrical Connections

A Comprehensive Survey

Harry A. Schafft

Electronic Technology Division

Institute for Applied Technology

/ . i National Bureau of Standards

Washington, D.C. 20234

-C.

,

e 4&3

+6 ^£^ J?&

U.S. DEPARTMENT OF COMMERCE, Peter G. Peterson, Secretary

Issued September 1972
National Bureau of Standards Technical Note 726

Nat. Bur. Stand. (U.S.), Tech. Note 726, 140 pages (Sept. 1972)

CODEN: NBTNAE

For sale by the Superintendent of Documents, U.S. Government Printing Office, Washington, D.C. 20402 (Order by SD Catalog No. C13.46:726). Price $1.25.
TABLE OF CONTENTS

Page

NOTATION ix

1. INTRODUCTION 2

2. GLOSSARY 6

3. WORD ABOUT UNITS 9

4. FABRICATION OF WIRE BONDS

4.1. Introduction 10

4.2. Wire 10

4.3. Bonding Surface 15

4.4. Gold Aluminum Interactions 19

4.5. Thermocompression Wire Bonds

4.5.1. Bonding Process 22

4.5.2. Bond Types 23

4.5.3. Conditions and Factors in Fabrication 27 4.6. Ultrasonic Wire Bonds

4.6.1. Bonding Process 31

4.6.2. Fabrication Procedure 35

4.6.3. Conditions and Factors in Fabrication 37 4.6.4. Factors Related to Bonding Equipment

4.6.4.1. Introduction 39

4.6.4.2. Ultrasonic Power Supply 40

4.6.4.3. Transducer and Coupler Assembly • . . 43

4.6.4.4. Bonding Tool 43

4.6.4.5. Mechanical Stage for Transducer-Coupler

Assembly 48

4.6.4.6. Chuck and Movable Stage for Device 49

4.6.4.7. Wire Clamps 49

5. METHODS FOR TESTING AND EVALUATING WIRE BONDS

5.1. Introduction 50

5.2. Visual Inspection

5.2.1. Description 52

5.2.2. Discussion 56

5.3. Pull Test

5.3.1. Description 57 5.3.2. Analysis 60

5.3.3. Discussion • 69

5.4. Shear Test

5.4.1. Description 73

5.4.2. Discussion 73

iii
Page

5.5. Air Blast Test

5.5.1. Description 74

5.5.2. Discussion 74 5.6. Push Test

5.6.1. Description 75

5.6.2. Discussion 75

5.7. Ultrasonic Stress Test 76

5.8. Centrifuge Test

5.8.1. Description 76

5.8.2. Analysis . 78 5.8.3. Discussion 80

5.9. Mechanical Shock Test

5.9.1. Description 81

5.9.2. Discussion 81

5.10. Variable Frequency Vibration Test

5.10.1. Description 84

5.10.2. Discussion 84

5.11. Vibration Fatigue Test

5.11.1. Description 85

5.11.2. Discussion 86

5.12. Short-Duration Stress Pulse Tests 86

5.13. Temperature Cycling and Thermal Shock

5.13.1. Description 87

5.13.2. Analysis 89

5.13.3. Discussion 93

5.14. Electrical Tests

5.14.1. Introduction 95

5.14.2. Bond Interface Resistance 95

5.14.3. Electrical Continuity 96

5.14.4. Noise Measurements 97

5.15. Ultrasonic Bond Monitoring

5.15.1. Description 97

5.15.2. Discussion 100

5.16. Correlation 101

6. SUMMARY 103

7

.

APPENDIXES

7.1. Appendix A. Estimate of Lowest Resonant Frequency of Wire Bonds 106

7.2. Appendix B. Expression Relating Four Measures for Pull Rate 110

iv
Page

8 . REFERENCES

8.1. Introduction 112

8.2. Citations 113

8.3. Addresses 123

8.4. Abbreviations 124

SUBJECT INDEX 126

LIST OF FIGURES

Figure 1. Sketch of a thermocompression ball-stitch wire bond

(enclosed by the dashed line) with the various elements

of the wire bond indicated where failure can occur 3

Figure 2. Sketch of a direct and an expanded contact with a goldwire ball-bond to aluminum metallization 21

Figure 3. Procedure for making a ball-stitch wire bond 25

Figure 4. Sketch of typical wedge (left) and eyelet (right) bonding

tools and the wedge and eyelet bonds made with uhese

tools 26

Figure 5. SEM photomicrographs (460 X) of bond adhesion (lift-off)

patterns for first bonds made at different time settings

and constant force and ultrasonic power settings 33

Figure 6a. SEM photomicrograph of the lift-off pattern of a normal

bond made under laboratory conditions 34

Figure 6b. SEM photomicrograph of the lift-off pattern of a partially

removed bond made under laboratory conditions 34

Figure 7. Simplified procedure for making an ultrasonic wire bond

between a package terminal and semiconductor die with a typical wedge-type tool 36

Figure 8. Schematic dependence of the pull strength of a single bond

on the ultrasonic power setting of the bonding machine 39

Figure 9. Sketch of an ultrasonic bonding machine with basic components and accessories labeled 40

Figure 10. Typical vibrational amplitudes along a long and a short

ultrasonic bonding tool at tool extensions recommended

by the manufacturer 44

Figure 11. Vibrational amplitude along a long ultrasonic bonding tool

for a tool extension of 0.350 in (0.889 cm), A, and a tool

extension of 0.335 in (0.851 cm) , B .

-. 46

Figure 12. Sketches of the bonding end of an ultrasonic bonding tool 47

Figure 13. Geometric variables for the double-bond pull test 61

Figure 14. Dependence of Fwt /F on 6 for various ratios 6 d to 6

fc 61
Page

Figure 15. F/Fw an<* *w/F as functions of d/h are graphed, for a singlelevel, double-bond pull test 64

Figure 16a. Fwt/F as a function of d/h for different values of H/h and

for a = 1/4 is graphed, for a two-level, double-bond pull

test 65

Figure 16b. Fwd^ as a function of d/h for different values of H/h and

for ex = 1/4 is graphed, for a two-level, double-bond pull

test 65

Figure 17. Dependence of AF(t) on pulling angle, <j>]_, for different values

of (1 - a)d/(h + H) and for (j) = 67

Figure 18. Dependence of AF(t) on the ratio H^/H for different values of

aH/h and for <f> = 67

Figure 19. Dependence of AF(t) and AF(d) on d^/d for <|> = 67

Figure 20. Dependence of AF(t) and AF(d) on hj/h for <j> = 68

Figure 21. Dependence of AF(t) on ai/a for a = 0.5, <j> = 0, and different

values of d/h and H/d 68

Figure 22. Dependence of AF(t) on a^/a for a = 0.25, <f> = 0, and different

values of d/h and H/d 68

Figure 23. Tensile force, Fw, in the wire adjacent to the bonds in a single-level, 1-mil diameter, gold wire bond due to a centrifugal force of 30^000 g's directed perpendicularly

away from the bonding surface is graphed as a function of

d for different values of d/h 79

Figure 24. Change in contact angle, ^ - ip , versus initial value of the

ratio of the wire-loop height to the wire span, h /d , for a differential expansion between the wire and the bonding surface, (gw - B s ) (T - T ), of .004 92

Figure Al. Variation of the frequency coefficients, C^, with subtending

angle, 0, and the ratio of bond separation to maximum wire

span height, d/h, for each of three vibration modes, identified by the sketched modes for = 180 deg 107

Figure A2. Dependence of the resonant frequency, f ]_, on bond separation,

d, for the lateral mode of vibration for 1-mil diameter gold

and aluminum wire bonds with circular shaped wire spans

having the ratios of d/h equal to 2 (semicircular), 5, and

infinity (straight wire) 109

Figure Bl. Illustration of Double-Bond Pull Test Ill

LIST OF TABLES

Table 1. Characteristics the Ball, Stitch, and Wedge Thermocompression

Bonds 24

Table 2. Methods for Attaining Thermocompression Bonding Temperature 29

Table 3. Test Methods and Brief Descriptions 51

vi
Page

Table 4. Visual Inspection Reject Criteria for Condition of Wire,

Condition of Bonded Wire, and Placement of Bond 54

Table 5. Pull Test Specifications and Conditions 58

Table 6. Conditions Used or Recommended in Centrifuge Test Methods 77

Table 7. Conditions Used or Recommended in Mechanical Shock Test

Methods 82

Table 8. Conditions Used or Recommended in Variable Frequency

Vibration Test 85

Table 9. Conditions Used or Recommended in Vibration Fatigue

Tests 86

Table 10. Conditions Used or Recommended in Temperature Cycling and

Thermal Shock Tests 88

vii

NOTATION*

d - horizontal bond separation: the horizontal distance between the two bonds of a wire bond. When used, the subscripts o and 1 denote an initial and an altered value, respectively.

E - modulus of elasticity (Young's modulus).

F - pull force applied by the pulling probe in a pull test or the pull

strength of the wire bond.

Fw

- tensile force in a wire.

F„j, F„t - tensile force in the wire on the die and on the terminal side of the

wire bond, respectively.

AF(d), AF(t) - ratio of the difference in the pull strength produced by changes in the

geometric parameters of the pull test from some initial set of values, to the initial pull strength for wire rupture on the die and terminal sides,

respectively.

g

- acceleration of gravity.

G - centrifugal acceleration (in units of gravity)

.

h - height of the apex of the wire loop above the terminal bonding surface.

When used, the subscripts o and 1 denote an initial and an altered value,

respectively.

H - height difference between the die and the terminal bonding surfaces.

When used, the subscripts o and 1 denote an initial and an altered value,

respectively.

k - ratio of the bending to the torsional stiffness (see footnote t on page

106)

.

TQ , T - initial and final ambient temperatures, respectively.

a - ratio of the horizontal distance from the bond on the terminal to where

the wire is pulled in a pull test, to the distance d. When used, the

subscripts o and 1 denote an initial and an altered value, respectively.

$w , 3s ~ thermal coefficients of linear expansion of the wire and the bonding

surface material, respectively.

a , 9t ~ contact angles: angles that the wire makes with the bonding surfaces of

the die and terminal, respectively.

p - wire density.

(j> - angle from the normal to the bonding surfaces and in the plane which

includes the two bonds, that the probe is pulled in the pull test.

iJ/ , ty - initial and final contact angles that the wire makes with the bonding

surface for a single-level wire bond, respectively, for a change in the

ambient temperature.

itNotations are not included where their definition and use are on the same page or in

the same tahle.

ix
ACKNOWLEDGMENT

It is my pleasure to acknowledge the help of and extend my sincere appreciation to the

many individuals who generously gave of their valuable time and experience during visits to several government laboratories and many development and production facilities of semiconductor device, bonding machine, and bonding tool manufacturing companies.

To those at the National Bureau of Standards, my thanks go specifically to Kathryn 0. Leedy, Herbert K. Kessler and especially to George G. Harman for many illuminating technical

discussions and suggestions regarding the contents of the paper; to Richard L. Mattis,

Willie E. Phillips, and Seymour Edelman who reviewed various parts of the paper and calculations, to Elaine C. W. Cohen for her help in collecting much of the literature; to Leo R. Williams and Edgar C. Watts for the drawings; to Terry A. Schultz, Ruth E. Joel, and Kaye E. Dodson for typing various drafts of the paper; to Mary Jane Ronas for her diligence in proofreading the final draft; and to Marilyn L. Stream who typed the final and camera-copy draft

with accuracy, with alertness to detail and consistency, and with good humor. Finally, I am especially grateful and indebted to W. Murray Bullis who, as the writing proceeded, made

many important technical- and editorial-related suggestions which were of great value to me

in improving the usefulness and clarity of the paper.
TESTING AND FABRICATION OF WIRE-BOND ELECTRICAL

CONNECTIONS — A COMPREHENSIVE SURVEY

Harry A. Schafft

Abstract

The fabrication and testing of wire-bond electrical connections used in integrated circuits, hybrid circuits, and low-power discrete semiconductor devices are

surveyed comprehensively. The survey is generally restricted to wire-bond electrical connections where the wire diameter is less than 2 mils and where the wire

is bonded either by thermocompressive or ultrasonic means. Under the general

heading of fabrication, the essential features of the thermocompression and ultrasonic bonding processes, the fabrication procedures, and the characteristics of

the constituent materials of the wire bond pertinent to high reliability are surveyed. Also included is a review of the interaction of gold and aluminum as one

of the primary failure mechanisms in wire bonds. Both new and old test methods are surveyed with emphasis on their capabilities and limitations. In particular,

the following test methods are discussed: visual inspection; pull, shear, air

blast, push, ultrasonic stress, centrifuge, mechanical shock, variable frequency

vibration, vibration fatigue, short-duration stress pulse, temperature cycling,

thermal shock, bond interface resistance, and electrical continuity tests; noise

measurement; and ultrasonic bond monitoring. Analyses of some of the methods

with regard to the stress that the test imposes on the wire bond have been made

and the results are used in discussing the relevant methods.

Key Words: Bonding; degradation (wire bond); discrete devices; electrical

interconnection; fabrication (wire bond); failure (wire bond); hybrid circuits;

integrated circuits; microelectronics; reliability; survey (wire bond); semiconductor devices; testing (wire bond); thermocompression bonding; ultrasonic

bonding; wire bond.
wire bonds' (importance of)

reliability

wire bond components

reliability

improvement

purpose

1. INTRODUCTION

The use of wire is still the predominant way of electrically connecting

the semiconductor dice and package terminals in integrated circuits, hybrid

circuits, and low-power discrete semiconductor devices.* The wire used is

typically 1 mil in diameter and may be as small as 0.7 mil. The reliability

of most of these devices is very high, but because of the large number of devices used in many present-day high-reliability systems, maintaining and increasing the component reliability remains a function of critical importance.

Because a significant fraction of the failures that do occur are failures of

the electrical connection [65B2], [68C1] , [68S2], [69H2], [69L1], [6901],

[70M3], the improvement and testing of these connections is a major concern. The term wire bond is used in this paper to refer to all the components

of this wire electrical connection. These components are illustrated in a sketch of a thermocompression ball-stitch wire bond in figure 1; they are the

wire, the metal bonding surfaces, and the adjacent underlying supportive

material. Each one of the elements indicated in the figure is a possible

failure point of the wire bond.

Most wire bonds do not fail. The problem is not that the technology is

unavailable to make highly reliable wire bonds. It is. Rather, the problem

is the inability to make the same wire bond every time and, the corollary, the

inability to identify sufficiently early those which will eventually fail during subsequent tests, in handling, or in use. Simply put, the two critical

areas for reliability improvement of wire bonds are the control of the manufacturing processes and the methods for testing and evaluating wire bonds.

The purpose of this paper § is to evaluate and review these two critical

areas for wire bonds with wire diameters less than 2 mils. Consequently there

*0ther connecting methods such as spider bonding, beam leads and other face-down schemes are

in growing use. Although setup costs are large, these all offer the advantage of being able

to make many bonds simultaneously. The more efficient bonding process and the reduction in

the human element in production offer the potential for lower cost and greater reliability.

Nevertheless, the vast experience with wire connections plus the inherent flexibility of

their use will continue to make this method important for some time in the future, in the

opinion of many experts [68P1], [71B2], [71J1], [71M1]. Uthe [71U1] has suggested that the use of wire bonds is, in fact, increasing. He reported that as a result of the economic re- versals in 1970-1971 many equipment engineering groups have been disbanded. As business begins to increase most companies no longer have the engineering talent and capability to implement the newer interconnecting methods. Thus, the problem of achieving higher reliability)

for wire bonds will continue to be of great significance. This is especially so for space

and military applications where extremely reliable devices are often required in relatively

small production lots that do not lend themselves to the use of multi-bonding schemes.

+Key words or phrases are placed in the margins to assist in scanning.

§This survey constitutes a portion of work on wire-bond evaluation which is being supported

by the National Bureau of Standards, the Defense Nuclear Agency, and the U. S. Navy

Strategic Systems Project Office (NAD, Crane) and is being performed at the National Bureau

of Standards as part of a comprehensive program on methods of measurement for semiconductor

materials, process control, and devices. Work in this program is reported in quarterly

reports and published as NBS Technical Notes [72B3].
METALLIZATION

OXIDE

METAL

FILM

Figure 1. Sketch of a thermocompression ball-stitch wire bond (enclosed by the dashed line)

with the various elements of the wire bond indicated where failure can occur. The elements

identified are as follows:

a. Wire.

b

.

Heel of the bond

.

c. Wire emerging out of the top of the wire material melted to form a ball prior to being

deformed in the process of making the ball bond.

d. Wire material melted to form a ball and deformed in the process of making the ball bond.

e. Wire material deformed in the process of making the bond and located over the bond between the wire and bonding surface.

f

.

Bond interface between the wire and the bonding surface.

g. Metal-film bonding surface. It may be multilayered.

h. Metal film (metallization) at the perimeter of the bond interface.

i. Interface between the metal-film bonding surface and an underlying insulating layer,

usually a silicon oxide,

j. Interface between the metal-film bonding surface and the terminal,

k. Interface between the insulating layer and the silicon substrate.

1. Silicon substrate.
approach

glossary

self consistency

subject index

references

information

sources

dearth of information

compensation for uncertainties

are two major sections to the paper. Section 4 addresses the fabrication of

wire bonds with emphasis on controlling the procedures and processes, and sec- tion 5 addresses both new and old test methods with emphasis on their capabilities and limitations. The various failure mechanisms and failure modes are discussed in the context of fabrication and testing in the appropriate

sections.

The paper has been written in a way which is intended to be useful both

to the relative novice and to those more experienced in wire-bond technology.

A glossary of terms used in the paper is provided primarily for the former who

may wish to proceed in order through the text. Separate sections of the paper are written in as self consistent a manner as possible to permit the more ex- perienced worker to read only those sections of interest to him. To further

assist the reader, a subject index is provided and each literature citation in

the list of references is followed by the page number (s) where the document is

cited in the text. Information included in this paper was obtained from the published literature; from unrestricted government reports;* and from interviews with engineers in government laboratories and in the development and production facilities of semiconductor device, bonding machine, and bonding tool manufacturing

companies

.

Review of the literature and the information obtained in personal visits

leads to the conclusion that there is a dearth of well documented experimental

work and a proliferation of hearsay and conclusions based on data from experiments where the test methods, the wire bonds tested, and the fabrication procedures have been incompletely specified. The survey paper reflects this paucity of definitive information. Incomplete information about optimum bonding

conditions and procedures abounds with regard to the condition of the bonding

surface, the wire material used, the bonding tool design, and the control re- quired of the bonding parameters. A similar situation exists for methods

used to evaluate wire bonds. All methods use one or more criteria to judge

the quality of the wire bond, but the relation of these criteria to the anticipated stresses and the potential reliability of the wire bond is at best

uncertain; test variables have not been explored sufficiently to determine

their effect on the value of the measure used; and, usually the methods are underspecif ied and the wire bond tested is insufficiently described.

In an attempt to compensate for the uncertainties in the stressed imposed

by the test and their relation to those stresses that the wire bond will need

to sustain in use, recourse is taken to making wire bonds as strong and as uniformly the same as possible and to testing to stress levels greater than

Some cleared information from limited distribution documents is included and referenced as private communications.
would appear necessary. For example, the centrifuge test is recommended by

many for testing wire bonds with gold wire but not for testing those with

aluminum wire; the stress imposed on the aluminum wire is significantly

less for the same acceleration level because of its lower density. But if

the reliability of the wire bond were primarily affected by a centrifugal

stress the method should be satisfactory for both wire materials.

A comprehensive bibliography on wire bonds by the author [72S1], which

bibliography

includes a detailed key word index, is a recommended companion paper to this

survey.
ball bond

bond

2. GLOSSARY

Bond formed with a capillary-type tool when the end of the wire has

first been formed into the shape of a ball (by melting the wire with,

for example, a flame as in a flame-off procedure). Also referred to as a nail-head bond. See figure 3.

The part of the wire bond that is associated with the volume of wire

deformed at an attachment point.

bond deformation — The change in the dimensions of the wire at an attachment point produced by the bonding tool in making the bond. The deformation is

usually measured in units of wire diameter.

bonding area

bonding schedule — The area within which the wire is to be attached to a terminal or die. — The values of the bonding variables used in bonding. For example; in

ultrasonic bonding, the values of the bonding force, time, and ultrasonic power.

bonding surface — The metallic surface or film to which the wire is or is to be interfaced and bonded.

bond lift-off

bond separation

capillary tool

The failure mode where the bonded wire separates from the bonding

surface.

The distance between the attachment points of the first and second

bonds of a wire bond.

A tool such that the wire is fed to the bonding surface of the tool

thru a bore located along the long axis of the tool.

direct contact — A contact such that the wire is bonded directly over the part of the

semiconductor die to be electrically connected, as opposed to an ex- panded contact. See figure 2.

elongation

expanded contact

eyelet tool

first bond

flame-off

flying-lead bond

foot length

gram-force

— The ratio of the increase in wire length at rupture in a tensile test

to the initial wire length, given in percent. — A contact such that the wire is bonded to an area remote from the part

of the semiconductor die to be electrically connected so that a lateral

interconnection path for the current is required. See figure 2. — A special wedge-type bonding tool designed to maximize the movement of

the wire surface in contact with the bonding surface while the wire is

deformed during the making of the bond. See figure 4. — The first bond in a sequence of two or more bonds made to form a wire

bond. — The procedure where the wire is severed by passing a flame across the

wire thereby melting it as in gold-wire thermocompression bonding to form a gold ball for making a ball bond. See figure 3. — See wire bond. — The long dimension of the bonding surface of a wedge-type bonding tool.

See figure 12. — A unit of force (nominally 9.8 mN) required to support a mass of one gram (1 gravity unit of acceleration * 1 gram of mass = 1 gram-force)

.

Colloquially, the term gram is used for the unit.
heel (of the bond) — The part of the wire that has been deformed by the heel of a wedge tool

used in making the bond (see figure 12) . The term is used primarily

with reference to ultrasonic wedge bonds

.

Kirkendall effect — The formation of voids, by diffusion across an interface between two materials, in the material having the greater diffusion rate into the

other.

Kirkendall voids

lift-off

loop (wire)

Voids formed by the Kirkendall effect.

See bond lift-off.

The curve formed by the wire between the attachment points at each end

of the wire.

loop height

overbonding

package

A measure of the deviation of the wire loop from the straight line between the attachment points of a wire bond. Usually, it is the maxi- mum perpendicular distance from this line to the wire loop.

metallization — The metal film (single or multilayered) on the semiconductor die used

to connect electrically different areas on the die.

Excessively deforming the wire, with the bonding tool, during the

bonding process.

The container for the semiconductor die or dice with terminals to provide electrical access to the inside of the container.

pad, bonding

peel, bond

post

Bonding area on the semiconductor die.

Similar to lift-off of the bond with the idea that the separation of

the wire from the bonding surface proceeds along the interface rather

than occurring all at once.

plastic device — A device where the package or the encapsulant for the semiconductor

die is made of such materials as epoxies, phenolics, silicones, etc.

search height

second bond

stitch bond

tail (of the bond)

tensile strength

terminal

— See terminal. — The ratio of the resonant frequency of the oscillating system to the

frequency band between half -power points (band width) . In the present

context, Q refers to the Q of the electro-mechanical system of an ultrasonic bonder, in particular to the sensitivity of the mechanical

resonance to changes in driving frequency. — The height of the bonding tool above the bonding area, at which final

adjustment in the location of the bonding area under the tool is made

prior to lowering the tool for bonding. — The second bond in a sequence of bonds made to form a wire bond. — A bond made with a capillary- type bonding tool when the wire is not

formed into a ball prior to bonding. — The free end of wire extending from the point where the wire is

attached to the bonding surface. — The force required to rupture a wire under tensile load. — A metal element used to provide electrical access to the inside of the

device package.
thermocompression

bonding

A process involving the use of pressure and temperature (not high

enough to cause melting) to join two materials.

tool, bonding — An instrument used to press the wire against the bonding surface in

the making of a bond.

ultrasonic bonding — A process involving the use of ultrasonic energy and pressure to join

two materials.

underbonding

wedge bond

wedge tool

wire

wire bond

Insufficiently deforming the wire, with the bonding tool, during the

banding process.

A bond made with a wedge tool. The term is usually used to differentiate thermocompression wedge bonds from ball and stitch bonds.

(Almost all ultrasonic bonds are wedge bonds.)

A bonding tool in the general form of a wedge with or without a wireguide hole to position the wire under the bonding face of the tool, as opposed to a capillary-type tool.

Unless otherwise specified, wire with a circular cross-section with a diameter less than 2 mils , and with either gold or aluminum as the

principal constituent material.

All the components of a wire electrical connection such as between

the terminal and the semiconductor die. These components are the

wire, the bonding surfaces, and the underlying supportive material.

See figure 1.
3. WORD ABOUT UNITS

Most of the data referenced in this paper that are not given in the

International System of units are followed in parentheses by the values in

the appropriate International System unit. General usage had dictated that

three exceptions to be made: (1) acceleration is given in units of gravity,

g (1 g = 9.8 m/s2 ), (2) the wire diameter is given in mils (1 mil - 25.4 ym)

,

and (3) the force exerted on the wire or wire bond is given in grams-force

(1 gf = 9.8 mN) . With respect to the last unit, common usage is not even to use the term grams-force but simply grams. In the interest of the proper

usage of units, the unit grams-force is used even though the data may have

been given in grams.
fabrication of wire bonds

4. FABRICATION OF WIRE BONDS

4.1. Introduction

wire

Au

Al

wire drawing

characteristics

(measurement)

Thermocompression and ultrasonic bonding are two principal processes

used in microelectronics to bond wire to metal surfaces on semiconductor di and on terminals. The essential features of the bonding process and fabrication procedures pertinent to high reliability are discussed separately according to bonding process. Preceding this is a discussion, in the context of

bonding and wire-bond reliability, of the wire and metal bonding surfaces

commonly used. Following this discussion is a review of one of the primary

wire-bond failure mechanisms, gold-aluminum interactions (intermetallic compound formation and Kirkendall voids)

.

4.2. Wire

Wires are usually made of either gold or aluminum. Gold wire is most

often used in thermocompression bonding while aluminum wire is most often

used in ultrasonic bonding. Wire hardened to some degree is needed to ease handling and aligning the fine wire while fabricating wire bonds. Gold tends

to age to its amorphous state with a consequent decrease in tensile strength.

Because hard-drawn gold wire ages significantly at room temperature, the

softer and relatively more stable stress-relieved wire is recommended [64C1].

Pure aluminum cannot be hardened sufficiently to allow it to be drawn to a diameter of 1 mil [69P1], [6902]. Therefore, aluminum is usually hardened by

adding about one percent of an impurity such as silicon or, less frequently,

magnesium.

Techniques other than those for standard wire drawing are being developed

to form small diameter wires. One is a continuous hydrostatic extrusion process [70S1]. Another is a process for drawing copper-cladded aluminum wire

[6902]; the copper is etched away after drawing to the desired size. While

pure aluminum may be drawn in these ways it may not be sufficiently hard to be handled conveniently.

A number of methods are available for measuring the different characteristics of the wire. ASTM standard methods are available for measuring wire

dimensions [70A4], [70A5] , tensile strength [70A6], and resistivity [70A7].

A technique which permits the automatic recording of load-strain curves has

been recently described by Hart [71H3] although at least one commercial tensile testing machine already incorporates such a feature. In an effort to determine the hardness uniformity of the wire in terms of some deformation measure more closely related to bonding conditions, the use of a wedge to be

applied with a given force along sections of the wire has been considered

10
[71B4], With the high intensity available from a laser it has become

possible to use simple diffraction effects to measure continuously, with

1/2 percent accuracy, the wire radius as it is drawn [66K2], [69G3]. Be- cause the intensity of the diffraction pattern is also a function of the

reflectivity of the surface the method can be used to monitor the condition

of the wire surface.

An ASTM Standard specification for gold wire is available [70A3] but

not for aluminum wire. This is not because the specifications of gold wire

are more important or critical, rather it is due to the fact that gold wire

was first used extensively. The ultrasonic bonding technology for using

aluminum wire lagged initially. Also, the problems with specifying alumi- num wire are more complex, as will be discussed later, and this too has

hindered standardization. A standard specification for hardened aluminum

wire is needed to define the pertinent wire characteristics, and how to measure them, so that wire of a consistent bonding quality may be produced,

purchased, and used.

The three most often specified wire parameters when ordering or describing wire are the diameter, tensile strength, and elongation. It is

important to have wire of a constant and known cross sectional dimension

because the bonding conditions, to be discussed later, depend on the mass

of wire involved in making both thermocompression and ultrasonic bonds.

The specification of the tensile strength and elongation is somewhat less

important for gold wire in thermocompression bonding than for aluminum

wire in ultrasonic bonding. For thermocompression bonding, changes introduced during the pre-heating and the heating during bonding are probably more important than the initial condition of the gold wire.

For ultrasonic bonding with aluminum wire, a low elongation is re- quired so that the wire may be broken after the second bond of the wire

bond is made without undue distortion of the adjacent wire which will be

used in the first bond of the next wire bond. If the elongation is too large, such distortions may be sufficient to cause the next bond to be

inferior. Too great an elongation may also result in an excessively long

tail in the second bond. The elongation that is usually specified is from

1 to 2 percent. Ravi and Philofsky [71R1] have pointed out that the measured elongation of aluminum wire depends on the strain rate and gage

length. Thus when specifying an elongation specification the method of

test is important.

Relatively hard wire is needed to minimize deformation of the wire

during ultrasonic bonding of small diameter (< 2 mil) wire. The range of

tensile strength used is from about 12 to 20 grams-force. Cox et at. [70C1]

standards

wire

specifications

for TC bonding

for US bonding

elongation

tensile strength

*Also S. Bonis and T. Salser, Raytheon Co., Sudbury, Mass. 01776; private communication.

11
handling

contaminat ion

degradation

Al + 1% Si wire

problems

recommend a range of 14 to 16 grams-force, saying that use of a higher tensil

strength (or harder) wire may fracture the silicon under the bond to the die.

The use of lower tensile strength (or softer) wire reduced the range of the

bonding schedules for which satisfactory bonds may be achieved. Also, the

softer wire is more difficult to handle and align under the bonding tool.

Care in handling the fine wire and protecting it from stress in all stepi

before bonding is important [67B2], [70A3], Wire is wound in a single layer on the spool to avoid binding and to aid in visual inspection of the wire.

Dropping the spool may dislodge the wire and interfere with free despooling

of the wire and the distortion due to the tensile stress on the wire in despooling can result in inferior wire bonds [68S1] . It is important to avoid

extremes in temperature to which the spooled wire is exposed, such as in

transport, because the difference in the thermal expansion coefficient of the

wire and spool may result in undue tensile stress or shifts of wire location

[64C1].It is vitally important that the wire be free of surface contamination.

A serious potential contamination problem is the insufficient removal of the

lubricating material used in drawing the wire. These lubricants are essen- tially transparent, and film residues a few hundredths of a micrometer thick on the wire are difficult to detect by optical or other means. In addition

to the possiblity of interf erring with making the bond, their presence can

lead to problems such as corrosion of aluminum metallization [71L1] or later

device degradation due to the water or ionic contaminants that may be included with the lubricant. All wire cleaning must be done by the manufacturer because the wire can neither be properly cleaned on the spool nor be

respooled.

Devaney [70D1] has shown striking SEM photos of wire bonds made with

gold wire contaminated with wire lubricants and with human contact.

Scarbrough and Auchterlonie [67S2] presented an example of how the transistor

gain can be degraded by organic residues of the lubricant on the aluminum

wire used. They postulated the possibility that some of the lubricant was

trapped below the surface of the wire during drawing. The problems that

alkali residues on magnesium-doped aluminum wire have caused will be discussed later in this section.

The addition of 1 percent silicon to harden aluminum for drawing and

bonding has resulted in numerous problems. These problems occur as a result

of the very low solubility of silicon in aluminum at room temperatures [67V1]

.

The instability of this metallurgical system leads to an annealing of the

wire in time, even at room temperature [69U1] . Hence such wire is considered

to have a limited shelf-life because the ease of handling and the adjustment

of the bonding variables depend on the hardness of the wire. In extreme

cases, silicon precipitates in localized regions causing gross non-uniformity

12
Si precipitates

tensile strength

in the hardness of the wire. When bonding with such wire, silicon precipitates may cause a fracture in the silicon under the bond of the die. Grain

growth can occur during exposure to the high temperatures (^ 500°C) employed

in sealing some ceramic packages. These grains can span the wire and severely reduce the ability of the wire to withstand mechanical stress [66K1],

[69P1], and [71R1]. Also, the rate of cooling from 500°C, where one percent

of silicon is soluble in aluminum, can affect the amount of silicon precipitated; the slower the cooling rate, the greater the precipitation.

A number of studies have been made of the mechanical and physical

properties of commercially available aluminum wire with 1 percent silicon

and the dependence of these properties on temperature and time [70C1],

[70P1], [71R1]. Large differences in the size and distribution of the

silicon precipitates in the aluminum wire as received from the manufacturer

have been reported [70C1], [71R1]. Ravi and Philofsky [71R1] found that

the most striking effect of these differences was upon the ductility of the ductility

wire. A maximum in the ductility was observed at a strain rate of 10-l+ s-1

for a uniform distribution of silicon particles in the tens-of-nanometers

size range. In wire with larger sized silicon particles, but still smaller

than 2 um in extent, this maximum in ductility was not observed. The ductility and the tensile strength was not otherwise changed by the increase

in particle size. However, for wire with silicon particles larger than

about 2 um the tensile strength was reduced. Cox et at. [70C1] found

significant differences in the distribution and size of silicon in wires

produced by different manufacturers to the same specifications. All

aluminum wires with silicon exhibit a reduction in tensile strength and

elongation with exposure to elevated temperatures as the silicon precipitates grow in size. Any initial differences between wires become small

after such aging, therefore high tensile strength wires show the more pronounced reductions. For example, the wire data from Prankatz and Collins

[70P1] shows about a 50 percent reduction in an initial tensile strength

of about 20 grams-force after about 100 hours storage at 150°C, or after

about 1 hour storage at 200°C. Fully annealed wire was obtained after

one hour storage at 300°C.

The use of wire in which the silicon is uniformly distributed with

individual silicon particles too small to be seen with a scanning electron microscope at a magnification of 3000 X has been recommended by Cox

et at. [70C1], This recommendation and a similar one by Ravi and

Philofsky [71R1] are not based on any reliability data; rather, they are

based more on an intuitive feeling for what the optimum conditions of

silicon dispersion should be. It is to be expected that the more finely

dispersed the silicon is, the longer it should take before significant

precipitation occurs.

13

effect of temperature

Si particle size
bonding schedule

stabilized wire

Al + 1% Mg wire

Of significance to bonding with wires having different mechanical properties is the observation of Cox et al. [70C1] that such differences affect the

bonding process. Thus if a different wire is used, a change in the bonding

schedule may be required. They also found that wire bonds made with a higher

tensile strength wire were less sensitive to changes in the bonding schedule

than those made with a lower tensile strength wire. The example they documented, however, compared a harder wire with silicon inclusions to a softer

wire with no silicon inclusions.* It would be interesting to compare the results using wires with the same tensile strength and different sized silicon

particles to evaluate better their recommendation for using wire with no "visible" inclusions.

Recently wire manufacturers have introduced a stabilized silicon-hardened

aluminum wire by adding a third element such as titanium to produce a ternary

alloy which is thermodynamically stable over the temperature range of interest.

A simpler alternative has been available. Magnesium has been substituted for

silicon to harden the aluminum wire. A 1 percent solid solution of magnesium

in aluminum is stable at room temperatures [67V1], Such wire appears to have

essentially the same mechanical and bonding characteristics as aluminum wire

with 1 percent silicon [69U1], [69P1], [70P1] with the added advantage that

the wire exhibits superior fatigue characteristics''" [72R1] and its tensile

strength is less affected by exposure to high temperature [70P1], However,

Plough et al. [69P1] and Davis [70D2] found that the presence of aluminum

wire with 1 percent magnesium caused transistor gain degradation after high

temperature storage. Uthe [69U1] implied the existence of such a problem.

These reports served to bring to the surface suspicions about the use of this

wire that others had expressed privately. However Pankratz and Collins [70P1]

have questioned the basis of these suspicions. Their studies showed that devices bonded with aluminum wire with 1 percent silicon degraded no less with

high temperature storage than those bonded with aluminum wire with 1 percent

magnesium. More recently in a continuation of work reported by Plough et al.

[69P1], evidence was uncovered to show that the earlier reported device degradation while using aluminum wire with 1 percent magnesium resulted from using

wire with an alkali-containing wire lubricant residue. When an alkali-free

wire-lubricant was used no such degradation was observed. § This development

gives added support to renewed consideration of the use of aluminum wire with

1 percent magnesium.

*The two wires had slightly different diameters. The tensile strengths, normalized to 1 mil

diameter wire, were 17.3 gf and 15.5 gf, respectively.

TAn important consideration for devices to be subjected to a large number of on-off power

cycles (see section 5.13.2).

»H. L. Floyd, Jr., Sandia Corp., Albuquerque, New Mexico 87115, private communication.

14
ribbon wire

metal film (die)

Kessler [69B5] has suggested the use of ribbon wire for making ultrasonic bonds instead of round wire. Preliminary work with ribbon wire with

essentially the same cross-sectional area as 1 mil diameter wire (1.5 x 0.5

mils) has shown that less pressure is required to make the bond. Thus, the

bond has a smaller deformation and the bond area is no greater than the area needed for round wire. In addition to the advantages of greater ease in

handling and aligning, the use of ribbon wire offers the opportunity of

strengthening the normally weakest part of wire bond, the heel of the first

bond. Because less pressure is required, higher tensile strength wire can be used without fracturing the silicon when bonding to the die. The resulting decrease in the necessary deformation appears to strengthen significantly

the heel of the first bond [72B2], Work in evaluating the advantages of

using ribbon wire has indicated that the control of the bonding schedule is

less critical for ribbon wire than for round wire, as judged from the results of pull tests [72B1],

4.3. Bonding Surface

The metal film most used on the semiconductor die is aluminum. When

gold is used, it is to avoid problems associated with the formation of goldaluminum intermetallic compounds. Because gold does not adhere well to sili- con dioxide and direct contact of gold to silicon is to be avoided, other

metals must be incorporated to form a multilayer metallization system.

Schnable and Keen [69S2] have described extensively the advantages and

Al vs. multilimitations of using aluminum metallization and concluded that it offers layer system

significant advantages over any other single or multilayer metallization

system that has been considered for integrated circuit applications. They

agreed with Selikson [69S3] that multilayer metallization systems suffer

in comparison not so much from inherent material limitations, but rather

from fabrication complexity where high fabrication costs and potentially

lower yield are the problems."'' Cunningham and Harper [67C1] in an earlier

paper presented arguments for the use of a gold-molybdenum metallization

system over the use of a pure aluminum metallization. Their arguments were based primarily on the greater metallurgical stability of the former

when gold wire is used in the wire bond. They however did not speak to the

fabrication-related problems associated with the greater complexity of the

system.

*Gold diffuses rapidly in silicon at high temperatures that are encountered in such procedures as thermocompression bonding, the sealing of some ceramic packages, and high temperature storage. The presence of gold reduces the minority carrier lifetime in silicon

which can lead to degradation of the electrical characteristics of the device.

+Problems of poor adhesion and the use of the shear test to detect them are discussed by

Gill and Workman [67G1].

15
metal film

(terminal)

Au

Au-plating

specification

Al

factors affecting bond ing

TC bonding

TC bonding

(thick film)

The base metal of the terminal is usually an iron-nickel-cobalt (FeNiCo)

alloy [70A8], designed to be thermally compatible with the glass seal used in

many device packages. The terminal, as part of the package supplied to the

semiconductor device manufacturer, is plated with gold, in most cases. Even

though the usage of gold plating is very common, the methods for specifying

and measuring the plating with respect to such properties as purity, organic

co-deposits, thickness, roughness, hardness, and porosity, are still inadequate, as has been discussed by Antler [69A3], because too little is understood about the properties of gold platings and how to measure them. To avoid

gold-aluminum interactions on the terminal when aluminum wire is used, the

terminal may be covered with aluminum. In rare cases, the terminal is bare

and aluminum wire is bonded directed to the alloy.

Bonding to the die and the terminal can be affected by many film-related

factors: surface smoothness, film hardness and thickness, film preparation,

and surface contamination, to name a few, but bonding technology is still not

developed to the point where any consensus exists about the optimum film con- ditions for fabricating a given type of wire. Only a little of this kind of

information exists in the literature with supporting experimental data. Some

of the comments that have been made in the literature and in interviews about

factors that may affect bonding wire to metal films are reviewed in the next

several paragraphs. Comments relevant to thermocompression bonding are con- sidered first.

Howell and Slemmons [64H2] indicated that for thermocompression bonding,

the uniformity > composition, and thickness of the metallization were important

and that, in particular, surface irregularities can prevent adequate diffusion

across the wire-metallization interface and hence interfere with making a bond

(see section 4.5.1). To insure adequate quality of the metallization for

bonding they proposed that one or more test areas should be made available

for bond pull tests. Hill [64H1] reported that by improving the uniformity

of the aluminum metallization thickness it appeared that the reliability of

the gold wire thermocompression bonds had been improved, but results were not

shown.

Budd [69B3] evaluated the use of a representative group of thick-film

metallizations for making gold wire thermocompression and aluminum wire ultrasonic bonds. Both the composition and firing conditions of the thick film had an effect on the pull strength of the wire bonds. Goldfarb [71G2] reported

that for ultrasonic bonds with aluminum and gold wires to four thick-film gold

materials studied, increased firing time tended to result in an increase in

the pull strength of the gold wire bonds and a slight decrease in the strength

of the aluminum wire bonds. No systematic dependence on the thick-films

examined was found.

16
Goldfarb also reported severe degradation of aluminum wire ultrasonic

bonds to thick-film gold materials with exposure to a sealing temperature

cycle which included exposure to temperatures greater than 350°C for 10

minutes. The degradation was indicated by increased electrical resistance

of the bond, decreased pull strength, and increased frequency of bond liftoff in pull tests. Although not mentioned, the failure mechanism appears

to be the formation of Kirkendall voids beneath the bonds. Thus, it would

appear that gold-aluminum interfaces on thick-film conductors are to be

avoided if the wire bond is to be exposed to high temperatures.

In the case of aluminum metallization, the temperature at which the

film is sintered, to achieve adequate adhesion [65R2] , appears to have an effect on ultrasonic bonding. Leedy et at. [70B3] found a sintering temperature dependence for aluminum wire ultrasonic bonds to 0.5-ym thick

aluminum films on oxidized silicon. Pull strengths of wire bonds on films

sintered at 500°C and 550°C were statistically similar and greater than

those on films sintered at 425 6 C and at 577°C.

The hardness of the aluminum metallization is said to be important.

It should be somewhat softer than the wire so that surface irregularities

may be easily smeared out to better conform to the wire [69P1] . Too soft a metallization may cause problems: Davis [70D2] reported that "excessive"

thickness results in a soft metallization to which it is difficult to bond.

Also, some concern has been expressed that the use of a soft aluminum

metallization could result in filaments of aluminum being extruded from

under the bonded region during bonding which if subsequently dislodged

could cause short-circuits.

The thickness of aluminum or gold metallization can have an effect on bondability and on subsequent reliability. There is generally no difficulty

encountered in bonding to films with a thickness within the typical range

used (0.7 to 1.2 ym) . Inconsistent bonding has been reported on films of

thickness less than 0.5 ym [67R1]. To avoid subsequent bond failure due to intermetallic compound growth and Kirkendall voids (see next section) at the interface, Philofsky [71P1] suggested that the thickness of the metal

film be minimized, consistent with good bonding and device design. This

suggestion applied both when bonding gold wire to aluminum on the semiconductor die and when bonding aluminum wire to gold plated terminals.

Kashiwabara and Hattori [69K3] found that subsequent failure of aluminum

wire bonds to gold plated terminals did not occur at storage temperatures

less than 350°C if the width of the wire-film bond interface area was

greater than 4 times the thickness of the gold plating. They reported that

the width of the actual wire contact was typically about 0.6 of the wire

deformation (as viewed from above) for a 3 ym thick gold film and about

0.8 for a 15 ym thick film.

US bonding

(thick film)

US bonding

Al sintering

temperature

Al hardness

Al, Au thickness

17
When the device is to be subjected to thermal or power cycling, wire

Al, Au thickness

flexing at the heel of the bond will occur. Philofsky [71P1] suggested that

under these circumstances the thickness of the aluminum metallization should

be less than one-sixth of the wire thickness at the heel of gold-wire wedge or stitch bonds to avoid the growth of brittle intermetallic compounds up into

this region and consequent fracture of the wire at the heel of the bond; for

the case of aluminum wire bonds to gold-plated terminals he suggested that the

thickness of the plating should be less than one-third the wire thickness at the heel of the bond. If either of these design recommendations is not feasible he provided maximum safe exposure times for the wire bonds to elevated

temperatures, based on the kinetics of transformation to the intermetallic

compunds

.

surface

roughness Excessive roughness of the bonding and substrate surfaces has been said

to influence the quality of ultrasonic bonds. Considering the constraints on

the thickness of the metallization, a substrate roughness of 30 yin. (0.76 ym) was considered highly undesirable and one less than about 3 yin. (0.076 um) was preferred* [67L1], [67R1] . Davis [70D2] recommended that the bonding

surface roughness must be such that the area of the bond be large compared

with the peak-to-peak variations in the surface. As an example, he suggested

the surface finish be from 4 to 8 yin rms (0.1 to 0.2 ym) for a 1-mil diameter wire and a 1-mil (25 ym) tool foot-length. Johannesen [71J1] mentioned

that wire deformation is affected by surface roughness, but the roughness and

degree of effect was undefined.

termi

S 1° Making bonds of consistent quality to the terminal (usually an ironnickel-cobalt alloy) is often more difficult to do than making consistent

bonds to the die. This is because package technology is far less sophisticated and far less in tune with the needs of wire bonding. This is particuAu plating

surface larly true for ultrasonic bonding. The finish of the gold plating is often

not sufficiently uniform or consistent to allow high quality bonds to be made.

Johannesen [71J1] cautioned that from his experience, gold platings that show

distinct grain structure or discoloration are unacceptable for bonding. Also,

the terminal is often so high that it is insufficiently rigid for ultrasonic

bonding^ and the bonding surface on the terminal is often insufficiently flat

or horizontal. The remedy is to grind down the terminal to the desired height

and flatness. The smooth finish thought to be necessary for good bonding may

be difficult to achieve when gold plating the ground-down terminal. Generally,

*In this instance and in some to follow, information was not provided to indicate if these

values for roughness are for peak-to-peak, rms, or other measures.

tThe terminal height should not be a resonant length with respect to the ultrasonic driving

frequency or its harmonics [68U5] . Davis [70D2] has suggested that for sufficient rigidity

the terminal height should be no more than one-half to one-third the diameter of the

terminal

.

18
large-sized grains of the alloy are exposed on grinding. The pre-plating

etch tends to etch preferentially at these grain boundaries and produce a rough surface. Two alternative techniques have been reported that can provide sufficiently smooth, gold-plated finishes. Both methods eliminate the

pre-plating etch step and replace it with a thorough cleaning procedure. In one method the terminals are stone lapped while in the other method the terminals are chemically polished to the desired height.

Contamination on the bonding surface should be avoided. For thermocompression bonding, it interferes with intimate contact and interdiffusion

of the wire and metal film and contributes to making poorer bonds [64H2]

,

[67K1] . The problem of contamination may be considered to be less for ultrasonic bonding because of the ultrasonic agitation. However, even if the

bonding schedule can be adjusted to break thru the barrier, such adjustments

generally involve increases in power, time, or force and as such can result

in increased deformation at the heel of the bond and hence reduce the pull

strength of wire bonds at bonding sites where the barrier is not as severe.

""

A number of contaminants capable of interfering with bonding have been mentioned in the literature. They are residues of chemicals used in the photoresist [69P1] and package plating operations [72H2], water spots, silicon

monoxide, silicon dust (from scribing), and aluminum oxide [66H1] . A radiotracer technique has been used to detect photoresist residues [69H1]

.

Antle [64A1], [66A1] has indicated the possible usefulness of a friction

technique to study the effects of surface contamination and of atmospheric

ambients, but the work reported dealt with wire metals not generally used in

microelectronics

.

With regard to contamination-induced bond degradation, Horsting [72H2]

found that impurities (greater than about the 0.1 percent level) in the gold

plating bath used to gold plate nickel-plated, iron-nickel-cobalt terminals

could cause bond failures at the terminals after high-temperature storage.

In particular, he reported bond failures of aluminum wire, ultrasonic bonds

after storage of about 1000 hours at 200°C.

contamination

(surface)

contamination

(plating bath)

4.4. Gold Aluminum Interactions 5

intermetallic

A typical wire bond has a gold-aluminum interface either at the die or at compounds

the terminal. Gold-aluminum intermetallic compounds form at this interface at

*H. L. Floyd, Jr., Sandia Corp., Albuquerque, New Mexico 87115, private communication.

+G. G. Harman, National Bureau of Standards, Washington, D. C. 20234, unpublished results.

§For further reading about the subject of gold-aluminum interactions, the paper by Selikson

[69S3] serves as a good, recent review. Earlier papers by Cunningham [65C5] and Blech and

Sello [66B3] are also recommended. The paper by Philofsky [70P2] provides an informative

study of the growth of Kirkendall voids and the different intermetallic compounds with

temperature.

19
;e

temperature a rate that increases with temperature. Above a temperature between about 125

and 150°C [67C1], [6901], the growth rate becomes significant with respect to long-term reliability of wire bonds. This is within the range of some operating environments, lower than the maximum temperature specified in many storage

and thermal cycling tests, much lower than the temperatures at which thermocompression, and perhaps even ultrasonic bonding takes place, and much lower

than temperatures used when hermetically sealing many ceramic packages. Thus

there is ample opportunity for such growth to occur. The compounds are formed by the diffusion of gold and aluminum across

Kirkendall voids

their interface. Gold has the greater diffusion rate and, as a result, will

leave behind vacancies on the gold side. This effect is named after

Kirkendall because of his studies of the relative diffusion rates in a copperzinc couple where such an effect was first observed [47S1], The process of

Kirkendall void formation can lead to two kinds of failure: a mechanical

stress-induced fracture along the locus of voids and an electrical opencircuit caused by the coalescence of voids.

plagues vs. Because the colorful gold-aluminum compounds were more noticeable than

voids

voids, the compounds were believed initially to be responsible for the failures. Thus the use of such terms as the purple, black, or white plagues were coined and many papers were written about which and in what amounts the

five different compounds (Au^Al, Au5Al2» AU2AI, AuAl, and AuAl£ [58H1])

appeared in the bond region and how failure ensued. The importance of the

void formation proposed by Cunningham [65C5] was gradually recognized but

continuing interest in the compounds themselves prompted Cunningham [67C4] to reiterate later that the exact identification of the compounds, and their

colors, relative strengths, etc., are essentially immaterial. The important

question is the extent of pore formation which depends on the time-temperature

history of the bond.

The kinds of reliability problems that result from gold-aluminum interactions depend on the wire bond type and whether a direct or an expanded contact is used (see figure 2)

.

electrical Electrical failure can occur when gold wire ball bonds are made to aluminum expanded contacts because of the formation of an annular Kirkendall

opening about the bond. This opening is difficult to see because of its

small dimensions and because it is often located beneath the outer extension

of the ball. The annular opening may be so small that a voltage of a volt or less (such as might be applied in an electrical continuity test) can cause an

arc to form across the opening and thereby "heal" the open-circuit (see

Voids have also been seen on the aluminum side of gold-aluminum interfaces. This has been

explained by postulating that the diffusion rates of gold and aluminum are modified when

diffusion occurs thru interposed intermetallic compounds [69S3].

20
DIRECT EXPANDED

BALL BOND

X^VT7 ' W/ {/('/////'/ Si 02

/ 'Si''

Figure 2. Sketch of a direct and an expanded contact with a gold-wire ball-bond to aluminum

metallization. The dashed lines indicate current paths. The sketch is not to scale.

section 5.14.3). The development of these voids at the perimeter of the

bond is accompanied by increases in the electrical resistance of the bond

with time that can be measured. The rate of increase in resistance with

exposure to elevated temperatures is larger for thinner aluminum metallizations [67K2], [67S5], [68A3], Normally, the bond adherence of these ball

bonds is unimpaired by intermetallic compounds which reach to the oxide

[65C5], [65H2], [65R1], [66B3], [68A3]. The intermetallic compounds adhere

well to silicon dioxide and though brittle can sustain a greater tensile

stress than either gold or aluminum [70P2]. Any slight reduction in pull

strength after exposure to elevated temperatures is usually attributed to annealing of the wire. If the thickness of the metallization is kept small,

void formation is apparently limited because of the small amount of material

available for diffusion [66B3].

Mechanical failure can occur when gold ball bonds are made to sufficiently thick aluminum films so that the supply of aluminum for reaction

with the gold ball is essentially unlimited. In this case the void formation

at the interface results in a mechanically fragile bond after high temperature

storage. Howell and Kanz [65H2] found an appreciable reduction in the pull

strengths of gold wires bonded to 6 urn thick aluminum films after 44 hours

of storage at 200°C. If thin metallization films are used, void formation

is apparently again limited. Blech and Sello [66B3] found essentially no change in the average pull strength of gold ball bonds to aluminum films

with thicknesses of from 0.05 to 1.0 um after storage at 200°C for as long

as 2000 hours.

Similar degradation can occur if an aluminum wire bond is made to a gold plated terminal where the gold plating is too thick [69K3]. In this

mechanical

failure

21
film thickness

AuAl2 metallization

thermocompression bonding

bonding

mechanisms

intimate contact

temperature

case an increase in the electrical resistance of the bond may be observed. In addition, if the bond deformation is excessive, be it an aluminum wire ultrasonic bond to a gold plated terminal or a gold wire wedge bond to an aluminum

film, the brittle intermetallic compounds may extend up into the heel of the

bond, and the wire bond becomes very fragile to any bending stress that is

imposed, for example, in the initial stage of the pull test or in temperature

cycling tests [71P1].

Thus, to minimize degradation effects due to gold-aluminum interactions one must avoid bonding to thick metal films and avoid excessive bond deformation. Guidelines by Philofsky [71P1] and Kashiwabara and Hattori [69K3] have

been given in section 4.3.

A patent has been granted for a metal film which inhibits the degradation

of gold wire bonds due to Kirkendall voids, by substituting a metal film of

the intermetallic compound, AuAl£, for aluminum [68T2], No report of its use

has been seen in the literature.

4.5. Thermocompression Wire Bonds

4.5.1, Bonding Process

Thermocompression bonding is a joining process involving the application

of pressure at an elevated temperature (not high enough to cause melting) for

a certain duration. There is still some disagreement as to how" and why the

bond occurs [69T2] . No definitive view of the mechanisms involved in the

joining process has been found in the literature.

Intimate contact between the two materials to be joined is an essential

requirement for making a strong thermocompression bond. This requires that

any intervening surface films such as oxides, water, and contaminants be dispersed or penetrated in the process of making the bond. Pressure, temperature, and some lateral spreading of the wire at the interface to disperse any

surface films are all presumed to be important factors in achieving this intimate contact. Diffusion is considered by many to be basic to the bonding

process [67K1], [69S1] , [69T2], The elevated temperature is thought to promote initimate contact with a minimum applied force by lowering the compressive yield strength of the materials involved and to accelerate diffusion

[69S1].

While it is generally agreed that the bonding temperature should be

lower than the lowest eutectic temperature of the metal system, Baker and

Bryan [65B1] also mentioned the importance of the bonding temperature being

below that at which dislocations may form or be displaced in silicon. In

this regard it is interesting to note that Goetzberger [61G1] has reported

the generation of dislocations in the emitter-base junction as a result of

22
yield point

ball bonds

bonding at 320°C. However, the possibility that excessive pressure could

also have generated these dislocations cannot be ruled out. Another requirement for a strong thermocompression bond is that the

wire have a sufficiently low yield point at the bonding temperature so that

when the joining pressure is removed the relief of the stresses will not

tend to restore the wire to its original shape and thereby weaken or rupture

the bond [66M2], [69S1].*

4.5.2. Bond Types

Characteristics of ball, stitch, and wedge thermocompression bonds are summarized in table 1. The most common thermocompression wire bond consists

of a ball bond to the die and a stitch bond to the terminal. The bonding

sequence for this wire bond is sketched in figure 3. Only gold wire is used to make thermocompression ball bonds. Aluminum

wire is not used because the oxide formation during flame-off prevents the

formation of an aluminum ball. Baker and Bryan [65B1] attempted to form

aluminum balls by a variety of heating methods, even in protective atmospheres, but were unsuccessful in obtaining balls of uniform size. More

recently, Kessler [70B2] has reported preliminary success in forming

aluminum balls by passing a jet of inert gas heated to about 1000°C across

the wire. His purpose was, however, to make aluminum balls for ultrasonic

bonding.

When stitch bonds are made at both ends of the wire, a mechanism is

required for cutting the wire and for making a right angle bend in the wire.

The bend is required so that when the wire is retracted up the capillary the

bent portion of the wire may be positioned against the capillary in preparation for the next bond. Scissor cut-off techniques which also bend the wireend as desired are described by KBllner [66K3] and by Rasimenoks et at. [67R4].

Here too, as in figure 3, the remaining upright wire from the second stitch

bond of the sequence must be removed as part of a separate operation unless

this step is incorporated in the machine operation. Helda and La Point [68H5]

describe a means for removing the upright wire after the last stitch bond is

made. The bonding tool is raised slightly and vibrated to metal-fatigue the

wire adjacent to the bond. While the tool is vibrated it is also moved horizontally away from the bond to break the wire there and also to bend the wire

extending from the bottom of the tool in a way which will facilitate the

making of the next stitch bond.

After making a stitch bond to the last bonding area a second stitch bond s^°"d stitch

is sometimes made before the wire is cut. This is done as a precautionary

stitch bonds

bond

Actually, Madland et at. [66M2] expressed this requirement in terms of modulus of elasticity and Slemmons [69S1] in terms of ductility of the bonded materials.

23
Table 1 - Characteristics the Ball, Stitch, and Wedge Thermocompression Bonds

Bond Type Characteristics

Ball 1. The bond is made with a capillary tool after the

end of the gold wire has been formed into a ball

as shown in figure 3.

2. Reputation for being the most uniformly rugged.

3. Orientation freedom — the second bonding site

does not need to be aligned with respect to the

bonding machine.

4. Relatively large bonding area is needed to accom- modate the ball (~ 2.5 times the wire diameter).

5. Strength of wire bond is limited by the tensile

strength of the wire just above the ball which

has been fully annealed during flame-off.

Stitch 1. The bond is made with a capillary tool. Figure

3 shows the making of a stitch bond after a ball

bond has been made.

2. Bonding area needed is less than for the ball

bond.

3. Stitch bonding sites must be aligned with the

direction of the wire feed and the bonding

machine to avoid twisting the wire adjacent to the bonds.

4. Wire is deformed at the bond so that a significant reduction occurs in the wire cross-section

adjacent to the bond.

Wedge 1. The bond is made with a wedge-shaped tool as indicated in figure 4; the wire must be aligned

separately.

2. Bonding area needed is smaller than that needed

for stitch and ball bonds. (Hence they are often

used in high-frequency devices and where the

bonding areas are closely spaced.)

3. Slower procedure because wire must be aligned

separately under the tool.

4. Bonding sites must be aligned with the direction of the wire feed and the bonding machine

to avoid twisting the wire adjacent to the bonds.

5. Wire can be deformed at the bond so that a significant reduction occurs in the wire cross- section adjacent to the bonds.

24
CAPILLARY

TOOL

7

CUT-OFF

TORCH

GOLD WIRE

7/ / / /

DIE

/ /

PRESSURE

AND

HEAT

TO FORM

©

BALL

BOND

ff

27

DIE DIE

PRESSURE

AND

HEAT

TO FORM

TERMINAL

©

/////

TERMINAL

STITCH

BOND

7777777

TERMINAL

Figure 3. Procedure for making a ball-stitch wire bond: (1) Gold wire is fed thru the

capillary type bonding tool; a small flame is passed across the wire below the bottom of the

capillary to melt the wire and, by surface tension, form a gold ball at the end of the wire.

(2) The wire is retracted so that the ball is held against the bottom of the capillary tool.

(3) The tool is lowered to the contacting surface where the gold ball is pressed against the

contacting surface and the interface is brought to the bonding temperature to form a bond.

(4) The tool is raised with the wire bonded to the surface. (5) The bonding surface on the

terminal is positioned beneath the bonding tool. (6) The tool is lowered as in step 3 to make a bond. This bond and any subsequent bonds made before the wire is cut again are called

stitch bonds. (7) After the last stitch bond is made, the capillary tool is raised as the

wire is paid out and a flame is moved across the wire to sever it and form a ball on each

end of the wire. —The bonder is then ready for the next bonding cycle and to make another

ball bond. The upright wire extending from the stitch bond is removed by manually pulling

the wire. The wire will rupture adjacent to the stitch bond because of the reduced wire

cross-section there. Grable and Patzer [69G2] have described a technique for avoiding this

separate procedure. After making the last stitch bond the capillary is raised slightly and

a wire clamp is activated and pulled backward to rupture the wire. The capillary is then

raised and a flame swept across the wire end to form a ball in preparation for the next

bonding cycle.

25
wedge bonds

eyelet bond

measure to reduce possible degradation of the bond by the process of pulling

the wire [71B2]. However, no clear evidence of such degradation has been

reported.

Wedge bonds are made with a wedge tool instead of the capillary tool

used for stitch bonds. Most wedge tools for thermocompression bonding do

not have a wire-guide hole, so an additional mechanism is required to align

the wire beneath the tool. Otherwise, the bonding procedures are similar to those for stitch bonding. Stitch and wedge bonds are usually made with gold

wire but aluminum wire may also be used. In the latter case, a hot inert

gas has been used both to heat the wire and bonding pad and to reduce the

oxidation of the aluminum wire. Johnson [68J1] has described the use of

such a hot-gas technique for thermocompression bonding. Hill and Wrench

[68H3] found that use of a pulse-heated capillary tool facilitates bonding

with aluminum wire.

As a result of their interesting studies of the plastic flow of metal

which occurs during the wire deformation in the bonding process, Baker and

Jones [66B4] have suggested an improved wedge bond which they called the

eyelet-bond. By using a specially designed tool, the lateral plastic flow

of the wire at the wire-metallization interface is increased without significantly reducing the wire cross-section in the process of making the bond.

Sketches of the tool and an eyelet bond are shown in figure 4 with a typical

wedge tool and bond. They reported eyelet bonds to be superior in strength

and in reproducibility to ordinary wedge and ball bonds of the same metal

combinations. The problem of the separate alignment of the wire under the

Figure 4. Sketch of typical wedge (left) and eyelet (right) bonding tools and the wedge and

eyelet bonds made with these tools. Capillaries to feed and align the wire are not shown.

26
bonding schedule

tool required for wedge bonding is aggravated in the eyelet bond because the

wire now must be aligned directly under the central projection on the face

of the bonding tool (see figure 4). The optimum diameter of this projection

was reported to be equal to the diameter of the wire used. This severe

alignment problem may be a reason for the lack of more recent reports of the use of eyelet bonding.

4.5.3. Conditions and Factors in Fabrication

The three primary conditions in thermocompression bonding are force,

temperature, and time. In optimizing these conditions to make a reliable

wire bond it is important to realize that not only are they interdependent

but they are affected by other conditions and factors. Minor changes in

these variables can cause significant differences in the bonding characteristics. Consequently, it is necessary to optimize experimentally the force,

temperature, and time. More than one combination of values for these three

conditions may produce a satisfactory bond. Of these combinations, Slemmons

[69S1] suggested that the most desirable bonding occurs with a combination

where the values for the force, temperature, and time are the lowest. It guidelines

may be assumed that this guideline includes such considerations as the

following: Short bonding time is desirable for production purposes. Low

bonding temperature is desirable to avoid degradation of the wire bonds

due to gold-aluminum interactions of the device due to alloying, and of

the passive components due to temperature-induced material decomposition

in hybrid devices. Low pressure is desirable to avoid fracturing or otherwise damaging the silicon beneath the bond. An additional guideline

can be recommended in the interest of reproducibility. The bonding

schedule should also be selected so that the sensitivity of the quality

of the wire bond to variations of these three conditions is minimized

which may not necessarily correspond to the above mentioned combination.

In general, the thermocompression bonding process is regarded as simpler and more forgiving that the ultrasonic bonding process, to be dis- bond and v;ire strength

cussed later. Some of the factors and parameters that have been mentioned

in the literature pertain to achieving a high joint strength between the

metal bonding surface and the wire. For most cases, however, it is not the

bond itself but rather the wire that is the weakest link in the wire bond.

In ball bonds the weakest link occurs in the high-temperature annealed wire

leading into the bond; in stitch and wedge bonds it occurs in the region of

the wire in which the cross-section has been reduced by the bonding tool.

A number of comprehensive review papers have been published which deal

reviews

with the various conditions and factors considered to be important in the

making of reliable thermocompression wire bonds [64H2], [65H2], [67K1].

27
force

temperature

time

ball size

Some of the highlights of these and other papers are included in the paragraphs that follow.

The bonding forces generally recommended are of the order of tens of

grams-force. Too large a force may damage the semiconductor substrate or excessively deform the wire. Often the force suggested is defined in terms

of the degree of deformation produced, of the order of 50 percent of the wire

thickness [64H2], [68K1]. Factors such as wire hardness and size, bonding

temperature, and tool design are therefore important. Budd [69B3] found that

the bonding strength of gold wire thermocompression bonds to a variety of

thick-film conductors was significantly less sensitive to variations in

applied force than bonds made to thin-film metallizations.

The typical range of bonding temperatures specified is from about 300

to about 350°C. The five different ways that have been used to bring the

interface temperature to the desired level during the bonding procedure are listed with comments in table 2. Howell and Slemmons [64H2] stressed the

importance of controlling the interface temperature, once selected. They

considered a variation greater than ± 5°C to be excessive for achieving uniform bonding characteristics. Such control may be difficult to achieve in

practice, however, and no data are available to indicate the sensitivity of

the quality of the wire bond to such variations in interface temperature.

Antle [64A1], [66A1] proposed an interesting method for selecting an optimum

interface temperature after which optimum force and bonding time could be

experimentally determined. He noted that, theoretically, bond strength

should be directly proportional to the coefficient of friction between the

materials to be bonded, measured at the bonding temperature. Thus, by measuring the variation of the coefficient of friction with temperature for

various metal systems and conditions, the optimum interface temperature and

how carefully it should be maintained may be determined. To support his

proposal, Antle showed a correspondence of a peak in the coefficient of friction and in the shear bond strength with temperatures for silver wire on gold

film. Further exploration of this relationship has not been reported in the

literature.

The range of bonding times generally used extends from a large fraction

of a second to several seconds depending on the magnitude of the bonding

temperature and applied force. The control of the time is more important

when the pulse heated capillary [68H3] is used than when the other heating

methods are used. Because the maximum interface temperature is obtained at

the end of the heating pulse, the bonding time in this technique effectively

controls the peak interface temperature for bonding.

One of the factors that must be considered when making ball bonds is

the variation in the size of the ball formed in the flame-off procedure.

Such variation in ball size should be minimized because, for the same applied

28
Table 2 - Methods for Attaining Thermocompression Bonding Temperature

Methods Comments

I. Heat DeviceHolding Stage

Inconvenient delay to bring device to desired bonding

temperature, especially for device packages with low

thermal conductivity. [67K1]

Long time at bonding temperature required for complex

devices with many wire bonds leads to excessive formation of Kirkendall voids and intermetallic compounds

when gold-aluminum interfaces are involved. Degradation of other components may also occur. [68H3]

Variations in bonding interface temperature from bond

to bond may occur, for example, due to:

1. excessive air currents

2. progressive heating of tool

3. voids in die-to-header solder or other factors

affecting heat transfer. [64H1], [67K1]

II. Heat Tool Allows more rapid bonding and less heating to device

than method I. [67K1]

Tool must be heated to a temperature higher than the

bonding temperature because the device and stage will

act as a heat sink.

At these higher temperatures (^ 350°C) both the capillary heaters and the tungsten carbide tools tend to wear out faster. [67K1] Also, gold wire at this tem- perature will tend to deform excessively while bonding.

[68H3]

III. Heat Stage

and Tool

The temperature of both the stage and tool can be lower

than in either of methods I or II with consequent re- ductions in the problems cited.

IV. Pulse Heat Tool

(one pulse per

bond)

Device exposed to bonding temperature for only a short

time (no thermal shock related problems encountered

with capillary tool or substrates used). [68H3]

Tool wear is much less than if method II is used, [68H3]

V. Heat Bonding

Interface by

Hot-Gas Jet

Pulsed heating must be adjusted for different substrates

(in most hybrid circuits the heating is adjusted to make

bonds to the substrate requiring the most heating).

[68H3]

Inert gas is used to bring bonding interface to bonding

temperature. The tool, substrate, or both may also be

heated. [68J1]

Method reported useful for bonding aluminum wire. Hot

inert gas serves dual function of heating and preventing

oxidation. [65R1]

29
flame-off torch

tool material

tool finish

tool wear

tool area and edges

wire alignment

wire spool

force, a ball larger or smaller than normal is subjected to insufficient or excessive deformation, respectively. A small ball may also become wedged in

the capillary and halt further bonding until the capillary is either cleared

or replaced. The size of the ball is affected by the speed at which the

flame-off torch passes the wire, the size and consistency of the flame, and

general ambient conditions in the work area [65H2], [67K1], [68H3]. Hill and

Wrench [68H3] recommended the use of a torch with a ruby orifice to improve

the consistency of the flame cone. Most capillary bonding tools are made either of tungsten carbide (with a cobalt binder) or of glass. Recently, tools of other materials such as titanium carbide, sapphire, and ceramics have become available with claims

of superior wearing properties to those of tungsten carbide. Other things

being equal, the bonding schedule is affected by using tools of different

materials at least to the extent that the thermal conductivities are different.

A smooth tool surface is considered important to prevent the wire from

sticking to the tool after bonding; a 0.5 yin. (0.025 ym) finish has been

suggested [64H2], The bonding surface of the tool wears with use. Alloying

of the gold in the wire with the cobalt binder in the tungsten carbide tools

leads to roughening and chipping of the bonding face of the tool. Titanium

carbide tools do not require a binder and hence may wear less quickly. Operating life of a bonding tool depends on the operating temperature history as well as general care and skill of the bonding operator. Capillary tools for

making ball bonds may also wear along the inner diameter of the capillary at the bonding end of the tool. In this case, the gold from the ball is extruded

up into the capillary tending to reduce the diameter of the wire leading into

the ball. Under these conditions bonding pressure on the ball is small and

non-uniform.

The tool contact area and the radii of curvature of the tool contact

surface edges should be such that bonds with adequate contact area may be

achieved without excessive deformation or abrupt changes in the wire crosssection. A radius of curvature from one to four times the wire radius is

usually recommended.

The importance of wire alignment in the making of consistent wedge bonds

so that the wedge makes contact with the wire in the same way each time was

pointed out by Howell and Slemmons [64H2] . They suggest that the location of

the wire-feed assembly be as close as possible to the tool so that maximum

support and control can be exerted. In addition, the wire spool should be

large to minimize curvature of the wire and it should have only a single

layer of wire to avoid distortion and cohesive bonding between successive

layers

.

30
Considerations for selecting a satisfactory range of loop heights are

as follows: Too small a loop height results in excessive bending of the

wire immediately above the ball bond. The wire may also be subjected to significant tensile and bending stress as a result of differential thermal

expansion of the wire and connected components (see section 5.13.3 for

further remarks on this topic). Too large a loop height makes the system

less rigid and subject to short-circuits with other elements of the device

and package as a result of accelerational stress.

loop heights

4.6. Ultrasonic Wire Bonds

4.6.1. Bonding Process

ultrasonic

bonding

intimate contact In the ultrasonic bonding process ultrasonic energy and pressure are

used to achieve intimate contact of the wire and metal in a way which results in a bond at their interface. Both the pressure and ultrasonic energy are transmitted by a bonding tool which presses the wire against the bonding

surface and is vibrated at a forced ultrasonic frequency, typically about

60 kHz. The applied force establishes and maintains adequate physical contact prior to and during the time that the bonding tool is energized. Prior

to the application of ultrasonic energy, the applied force deforms the wire

somewhat. The actual area of contact is much smaller than the apparent area

of contact because of the sub-microscopic surface roughness which exists.

With the application of ultrasonic energy the temperature at the interface

rises, the wire is deformed further, and the interfacing surfaces are

smoothed, increasing the area of contact. Surface oxides and any other

surface films are disrupted and dispersed, and nascent metal surfaces are

thus brought into contact in a way which results in a bond.

The plasticity of the metal during the process results not only from

the temperature rise at the interface region but also from the ultrasonic

energy as well [68B1], For aluminum the application of sufficient ultrasonic power density has the effect of reducing the stress necessary for

plastic deformation during exposure and hardening aluminum after exposure;

increasing the temperature while the ultrasonic energy is applied decreases

the hardening effect [66L3].

The mechanisms involved in the actual bonding process are not fully

understood [65J1] , [68B1], [69P1] , [71J2]. The results of early and extensive studies of ultrasonic bonding by Weare et al. [60W1] suggested an interface melt hypothesis in which bonding was due to the formation of a thin melt hypothesis

molten film at the interface, caused by the dissipation of ultrasonic energy

there. Bikerman [66B6] asserted that this was the only way an adhesive bond

could be made. More recently, Uthe [69U1] has been a proponent of the existence of an interface melt, however thin it may be. Thus, the ultrasonic

31

wire plasticity

US softening and hardening

mechanisms
disagreement

temperature measurements

metallographic

exam

weld formation

bond could be called a "hot pressure bond" because of the pressure and high

interface temperatures involved.

Others have disagreed with the interface melt hypothesis and instead

have proposed that the process is much more complex [62P1], [65J1], [67J1],

[71J2]. Peterson et at. [62P1] and later Jones [67J1] made temperature measurements which led them to conclude that the maximum interface temperature

achieved was between 30 and 50 percent of the melting temperature of the

metal having the smaller melting temperature. Joshi [71J2] measured bond

interface temperatures while bonding wires of gold, aluminum, and copper to gold-plated copper-cons tantan thermocouple beads ground to the shape of a small thin disc and rigidly held by embedding in epoxy. He reported temperature increases of less than 70°C during bonding. When a deliberate decrease

in the bonding force was made, to promote slippage at the bonding interface

and thereby increase frictional heating, a temperature increase of 120°C was measured when bonding gold wire. Jones et at. [65J1] stated directly that no evidence for a melted zone has been seen and further that heat by itself does

not appear to play a significant role in the bonding process. Joshi [71J2]

has concurred. Supportive evidence to indicate that achieving a high temperature during bonding is not important has been provided recently by Joshi [71J2]

and Harman and Leedy [72H1] who have made gold-to-gold and aluminum-to-aluminum

ultrasonic wire bonds, respectively, at liquid nitrogen temperatures. Metallographic examinations of bond cross -sections of a variety of similar and dissimilar metals have been reported [62P1], [ 65Jl] » [71J2] . These examinations

have revealed a number of interfacial effects, such as interpenetration, and

surface film disruption and dispersion; working effects, such as plastic flow,

grain distortion, and edge extrusion; and heat effects, such as recrystallization, precipitation, phase transformation and diffusion; but no unifying concept of the ultrasonic bonding process has been formulated.

The initiation and growth of the welded area appears to occur at distinct

localized regions as described by Jones [67J1] rather than in a generally continuous way from the center of the contact area outward to the edge of the

wire contact, as postulated by Uthe [68U2-4], Scanning electron microscope

observations of welded areas as the bonding power and time is increased have

been reported by Leedy [69B3], [70B8], Bonding was initiated at discrete

spots generally at the heel or toe of the contact area (under the back or front of the bonding tool). Initiation points appeared randomly in these areas from bond to bond. With greater power and time these discrete bonding

sites expanded, and grew along the sides. Bonding at the center generally

did not occur. A development of such bonding sites is shown in the sequence

pictured in figure 5. In figure 6, a SEM photomicrograph of a lift-off pattern and one of a partially lifted bond illustrate the lack of bonding to the

central region. Silicon precipitates in the aluminum wire were postulated as

32
\.

-V

*—

s

J

Figure 5. SEM photomicrographs (460 X) of bond adhesion (lift-off) patterns for first bonds

made at different time settings and constant force and ultrasonic power settings. The time

is longest for case a and is decreased to the shortest value for case d. Settings used in case a result in a bond which will adhere to the pad; the shorter time settings do not produce an adherent bond. The surface of the bonding pad under the wire is smoothened by the

lateral movement of the wire produced when the tool is brought down on the wire with the

selected force and before ultrasonic power is applied. Note the apparent movement during

bonding exhibited in case d. This is thought to result from mechanical vibration in the

bonding machine. (From Leedy [70B3])

33
Figure 6A. SEM photomicrograph of the

lift-off pattern of a normal bond made

under laboratory conditions. The pattern, which clearly reveals the unwelded

center portion of the bonded region, was

exposed by carefully peeling back the

bond. Magnification: 510 X. (From

Leedy [70B8])

Figure 6B. SEM photomicrograph of the

lift-off pattern of a partially removed

bond made under laboratory conditions.

A similar pattern can be observed on both the wire and the pad. Again the

center portion of the bonded region is

not welded. Magnification: 900 X. (From Leedy [70B8])

intimate contact

the nucleating centers for the initiation points, hut preliminary observations with magnesium-doped aluminum wire, in which magnesium does not precipitate, showed similar bond growth patterns. Recent investigations by

Kessler [72B2] of lift-off patterns of ultrasonically bonded aluminum ribbon

wire, have indicated that the way welding proceeds spatially may be dependent

on the bonding schedule: for a high-power, short-time bonding schedule, primary welding occurred at the perimeter of the wire; while for a low-power,

long-time bonding schedule, primary welding occurred in the central region

as well as at the perimeter of the wire.

Aluminum wire is commonly used in making ultrasonic wire bonds. Uthe

[69U1] cited two factors responsible for the relative ease with which

aluminum wire ultrasonic bonds could be made: (1) the abrasive action of

the aluminum oxide (from the oxide coating) in an oxide-aluminum slurry

cleans and extends the bonding area and (2) the tensile strength of aluminum

decreases with temperature so that the soft, conforming aluminum adjacent to the interface can be well supported by the outer and cooler parts of the

wire. Contradicting the suggested importance of these two factors are the

generally non-continuous way that bonding proceeds, which does not easily

34
fit with the concept of an oxide-aluminum slurry, and the relative unimportance of temperature in bonding. The softening needed to achieve intimate

contact appears to be produced by absorption of ultrasonic energy [71J2],

[72H1]. Furthermore, gold which does not have an appreciable oxide nor the

strength-temperature characteristics of aluminum can also be bonded ultrasonically [65J1] , [68S1], [71G2], [71J2]. While no detailed studies have

been reported which have examined the relative ease of making aluminum and

gold wire ultrasonic bonds, Joshi [71J2] has at least indicated no significant differences in their bondability.

Ultrasonic bonding is a more critical process than thermocompression

bonding. The principal reasons are that the ultrasonic energy transmitted

to the wire and the bonding surfaces must be inferred and there is no clear

understanding of how this energy and its rate of transmission is related to making a satisfactory wire bond. Furthermore, a number of conditions and

factors (discussed in section 4.6.3) mutually affect the energy transmitted

to make the bond in ways that are usually not clear. Nevertheless, with

sufficient control of the empirically desired fabrication procedures highly

reliable wire bonds can be made. Two important advantages that ultrasonic

bonding has over thermocompression bonding are that it is fast and that no heating is required [67K1], [68H3] , [69S1J.

The process as normally used with a wedge tool which is lowered and

raised by pivoting, has three limitations that become particularly acute in

hybrid circuits [68H3] : (1) the bond surfaces must be uniformly rigid,

(2) the height of the die and terminal bonding surfaces cannot vary greatly

from device to device, and (3) the device must be rotated to align the two bonding sites with the wire. The first requirement relates to the need to couple the ultrasonic energy into the bonding area. The other two requirements are relaxed if ultrasonic ball-stitch bonding with gold wire is used

because the tool moves vertically to and from the bonding surface and there

is no need to align the two bonding sites to avoid twisting the wire in

moving from the first to the second bonding site. Consequently, gold-wire

ultrasonic bonding is finding more use in hybrid circuits [70E1].

US vs. TC bonding

limitations

Au-wire US bonding

4.6.2. Fabrication Procedure

A simplified procedure for making an ultrasonic wire bond between the

terminal and the semiconductor die of a device with a typical wedge tool is

sketched in figure 7. Usually there is insufficient room for the tool if

the first bond is attempted on the die. Hence the first bond is usually

made to the terminal as shown. The procedure for making ultrasonic bonds

figure 7

*Such rotation is inconvenient and it can contribute to operator error in devices with com- plex bonding pad location patterns [71B2].

35
WIRE

PRESSURE

AND

ULTRASONIC

ENERGY

TO FORM

V

TERMINAL TERMINAL

BONO

TERMINAL

TERMINAL

DIE

PRESSURE

AND

ULTRASONIC

ENERGY

TO BOND

Figure 7. Simplified procedure for making an ultrasonic wire bond between a package ter- minal and semiconductor die with a typical wedge-type tool. (1) Wire is located between the

bonding surface of the tool and the terminal bonding area. The bonding tool is then lowered

to its "first search" position (75 to 125 um above the bonding surface). This position is

preselected and determined by the average height of the surfaces to which the first bond is to be made. (2) The tool is lowered to press the wire against the bonding surface with a predetermined force. Ultrasonic energy is applied for a preset time to make the first bond.

(3) The tool is raised while the wire is paid out from the spool of wire (not shown). (4)

The bonding stage holding the device moves the second bonding site under the tool and the

tool is lowered to its "second search" position (similar to steps described in 1). (5) The

tool is lowered to make the second bond as in 2. (6) After the second bond has been made,

a wire clamp to the rear of the tool (not shown) closes and pulls back on the wire to break

the wire at the heel of the bond. The tool is then raised and the end of the wire is fed

out underneath the tool until the end is located somewhat beyond the front of the tool, as shown in 1. The bonder is ready to be located over the next first bond site to repeat the

bonding cycle.

36
with a capillary tool is similar except that means for bending the wire under

the tool for the first bond are needed. If gold-ball ultrasonic bonds are to be made, the additional step of forming the ball is required. This has been

described in the thermocompression bonding section.

Prior to the procedure in making a wire bond with a wedge tool, such as sketched in figure 7, an alignment procedure must be performed. The stage

holding the device must be manipulated by the operator to align the two bonding sites so that they lie under and in line with the direction of the

wire. This is done to avoid twisting the wire at the heel of the first bond

while moving the second bonding site beneath the bonding tool.

alignment

4.6.3. Conditions and Factors in Fabrication

The three primary conditions in ultrasonic bonding are the force, time,

and ultrasonic power. The ultrasonic power available to make the bond is

dependent on not only the power setting of the oscillator power supply but

also the frequency adjustment and other more subtle variables to be discussed

later. The force used in bonding must be large enough both to hold the wire

in place without slipping and to couple the ultrasonic energy into the bonding

locale without producing excessive deformation of the wire [67J1], It is

generally of the order of tens of grams-force. The specific value selected

depends on a number of factors, some of which are the size and design of the

bonding tool face, the size and hardness of the wire, and the sensitivity of

the substrate and any adjacent active areas to the bonding pressure. Order

of magnitude values for the input power to the transducer and the bonding

time are a few watts and tens to hundreds of milliseconds. A preference has

been expressed for a high-power, short-time to a low-power, long-time com- bination to avoid metal fatigue and the initiation of internal cracks [65J1]

,

[68U4], On the other hand a lower power and longer time schedule than

usually used has produced bonds with a better cosmetic appearance and a greater pull strength when ribbon wire was used [72B2]

.

No definitive values for the parameters of force, time, and power can

be given for an optimum bonding schedule. Only general recommendations can

be given which may be used in adjusting a given bonding machine for particular wire, bonding surfaces, semiconductor die, and package characteristics.

This is because many subtle factors can affect the quality of the wire bond.

Also, the uniqueness of each bonding machine must be considered in developing

the optimum bonding schedule because of the slight differences in friction,

looseness, rigidity and other mechanical tolerances [69U1].

The general guiding philosophy for establishing the optimum bonding

schedule is that the bonding parameters be adjusted so that it is the reproducibility of the process rather than the strength of the wire bond that is

maximized. That the strength of the wire bond may need to be compromised to

US power

force

time

schedule

bonding schedule

schedule

guidelines

37
figure 8

d

achieve greater control over the process [67R1] , [69K1], [70B8] , can be seen from the following remarks. The primary criterion for strength in optimizing

the bonding schedule is the pull strength of the wire bond as measured in a pull test (see section 5.3). There are two primary failure modes in the pull

test: rupture of the wire adjacent to the bond, where it has been deformed

and reduced in cross-sectional area, and rupture or lift-off of the bond at the interface between the wire and the metal film. A useful schematic

representation is shown in figure 8 where the rupture strength of the bond

and the tensile strength of the deformed wire adjacent to the bond are plotted

as a function of ultrasonic power, after Kashiwabara et at. [69K1], The

sketch shows with the dashed and dotted lines that as the power increases,

the tensile strength of the deformed wire decreases, indicating an increase

in wire deformation, while at the same time the adhesion strength of the bond

itself increases rapidly above some level of ultrasonic power. The failure

mode and the pull strength is determined by which of the two broken curves is

lower. The shaded stripe indicates the expected dependence of the pull

strength on the power setting for a single-bond pull test. The cross-over

point, at P , is altered by the time and force settings of the machine as well

as other factors (to be discussed later). For example, when the force and

time are reduced, P is increased without any reduction in the maximum pull

strength [69K1]. Uncontrolled small variations in the bonding conditions will

cause larger variations in the pull strength in the region where bond failures are expected because of the steepness of the slope of the bond strength curve,

as shown in figure 8. The width of the shaded stripe is meant to indicate the

variation in pull strength for a given power setting. Thus, if in this case

the ultrasonic power is selected to achieve a maximum pull strength, a larger

variation in pull strengths will be expected. This indeed has been found to be the case and has been the basis for the suggestion that the bonding schedule not be adjusted for maximum pull strength [69K1] , [70B8] because of the

penalty of a greater variation in pull strength,

optimizing

schedule The dependence of pull strength on ultrasonic power indicated by the

shaded stripe in figure 8 is similar to the dependence reported by Leedy

pull strength [70B8] of pull strength* with not only power, but with bonding time or with

bonding force. Similar curves of such dependencies have appeared elsewhere

[69P1], [70D2], [70P1]-, but for double-bond pull tests in which failure was allowed at either end of the wire bond. In these curves the pull-strength

maximum is sometimes more centrally located. The use of such curves to achieve an optimum bonding schedule has been described by Plough et at. [69P1], Davis [70D2], and Leedy [71B5]

.

Bonding machine settings were adjusted so that failures occurred only at the first bond of

the single-level wire bonds tested.

38
BOND STRENGTH

TENSILE STRENGTH OF DEFORMED WIRE

ULTRASONIC POWER

Figure 8. Schematic dependence of the pull strength of a single bond on the ultrasonic

power setting of the bonding machine. This is drawn from the schematic dependencies on the

ultrasonic power of the adhesion of the bond and the tensile strength of the heel of the

bond. The width of the shaded curve is meant to indicate the expected variation in pull

strength because of uncontrolled variations in the bonding process, and other variables.

The critical power setting, Pc

, determined by the crossover of the dotted and dashed curves indicates a broad dividing line below which lift-off failure and above which wire failure

is expected in a pull test. Fg represents the tensile strength of the undeformed wire.

The quality of the wire bond is also judged by wire deformation at and near the bond as well as by other features. This is discussed in sections

5.2 and 5.16.

4.6.4. Factors Related to Bonding Equipment

4.6.4.1. Introduction

Many of the factors that can affect bonding are associated with the

bonding machine. The basic components of the ultrasonic bonding machine

and its accessories are sketched in figure 9. Basic requirements for making

consistently high-quality wire bonds are: the precision control of the mechanical movements, the rigidity of the mechanical components and the bonding

requirements for process control

39
VIP—^

BONDING

TOOL

DEVICE

CHUCK

WIRE SPOOL

COUPLER

^

ASSEMBLY

PIVOT

COUNTER WEIGH'

TRANSDUCER

TO POWER SUPPLY

Figure 9. Sketch of an ultrasonic bonding machine with basic components and accessories

labeled.

resonance

surface, and the absence of mechanical vibrations [69U1], [70D2], [71H1],

[71J1]. The application of capacitor microphones and magnetic pickups for

diagnosing both mechanical and electrical problems in wire bonding equipment

have been described recently by Harman and Kessler [71H1]. They were able,

for example, to detect excessive relative movements between the tool and

bonding surface which could be traced to the cantilever type supports for the

transducer-coupler assembly housing, operation of the cam motor, and vibration

or shock transmitted through the floor or base. Uthe [69U1] has also suggested tests for diagnosing bonding problems, many of which are related to the bonding equipment.

4.6.4.2. Ultrasonic Power Supply

The ultrasonic power supply generally has two channels to provide different power levels and bonding times for bonding to the semiconductor die and

terminal bonding surfaces, respectively. The ultrasonic frequency is adjustable about a mean frequency of approximately 60 kHz to allow tuning the system to resonance, where the transmission of ultrasonic energy to the bonding

40
area via the bonding tool is most efficient. The system is designed to deliver significant energy only within a limited frequency range about

this mean frequency. The extent of this frequency range is determined by

the Q of the electro-mechanical system.

A preliminary step in adjusting the bonding machine is to tune the

power supply to achieve a resonance of the mechanical system driving the

bonding tool. In some power supplies the procedure recommended by the

manufacturer is to tune to electronic resonance. However, the interaction

of the electrical and mechanical systems is complex, and Harman [69B6] has

shown that the conditions for electrical and mechanical resonances may not

necessarily coincide. To avoid this and other problems Harman and Kessler

[71H1] have described procedures for tuning the bonding machine and reproducing a specific ultrasonic vibrational amplitude at the end of the bonding

tool. Control and reproducibility of the tool tip vibrational amplitude

(typically 1 to 2 urn, peak to peak) is considered to be important in making

consistent wire bonds but as yet no experiments have been completed in which

tool tip amplitude has been correlated with such quality measures as pull

strength or degree and appearance of deformation. However, after these

procedures were incorporated by one semiconductor device manufacturer, a reduction by more than a factor of two was achieved in the rejection rate

due to faulty bonds at visual inspection [71B5].

The Q of the mechanical system is an important consideration in the

control of the bonding process. Normally, a broad or low-Q resonance is

preferred so that drift of the oscillator frequency and changes in the

temperature or loading do not significantly detune the mechanical system

and reduce the ultrasonic power available to make the bond with a machine

in the fixed -frequency, uncompensated mode. Kessler [70B3], has described

a simple means of displaying the resonance curve of the tool displacement

as a function of frequency on an oscilloscope screen. The Q and the resonant frequency have been found to be dependent on

the output impedance of the power supply [70B3]. Measurements on the same transducer-coupler assembly and tool showed that the mechanical Q was

lowest (Q = 70) for a power supply with a constant current source, intermediate (Q ~ 135) with a constant power source, and highest (Q ~ 200) with

a constant voltage source. The use of these different power supplies produced a maximum change in the resonant frequency of about 350 Hz. Both resonant frequency and tool tip vibrational amplitude have been

found to be dependent on temperature [70B2]. Temperature changes of five

or more degrees Celsius may produce measurable changes in the operating

characteristics of the bonder. For example, if high-intensity wide-angle

lamps are used to illuminate the work area then appreciable heating

(5 to 10°C) of the transducer-coupler assembly normally occurs. Thus

tuning to resonance

system Q

power supply

impedance

temperature

effects

41
oscillator drift

frequency sweep

problems with

frequency sweep

slow sweep

feedback

u

such bonding machines should be tuned after the lamps have been turned on for a time long enough to establish thermal equilibrium.

To compensate for drift in the oscillator frequency and other changes

that detune the mechanical resonance, a swept-frequency mode of operation has

been used. Here the oscillator frequency is cycled repeatedly over a range

which hopefully includes the mechanical resonance frequency. Feelings about

this approach are mixed and no definitive evidence pro or con has been developed. A potential problem is that if the exciting frequency is sufficiently

different from the mechanical resonant frequency, other vibration modes may

be excited in which the tool tip moves in directions other than parallel with

the wire and bonding surface. This may occur in some cases if the driving

frequency is as little as 400 Hz different from the mechanical resonant

frequency.

The primary advantage of using the swept-frequency mode of operation is that bonding should be relatively independent of the mechanical resonant

frequency. However, Harman and Kessler [69B6] , [70B3] have questioned this

independence by showing that the variation of the tool tip vibration amplitude during the cycle is dependent on both the location of the mechanical

resonant frequency within the range swept and the way the frequency is modulated because of the lag in the response of the mechanical system to rapid

changes in the exciting frequency. It remains to be seen, however, whether

the differences they noted are sufficient to result in significant changes

in the quality of the wire bonds so fabricated.

A better case may be made for a single or several slow sweeps through resonance to make the bond [70B3]. If the frequency is modulated linearly

and slowly enough to allow the mechanical system to respond, the dynamic

response of the tip vibration amplitude envelope is similar to the static resonance curve. Under these conditions the initial coupling between the

tool and wire, and the completion of the bond are accomplished under lowpower conditions. Some believe that it is better to use a low power initially to avoid heating at the tool-wire interface because of possible slippage

there, and to use low power at the end of the power cycle to reduce the possibility of inducing wire fatigue. Jones [67J1] has recommended the use of

this kind of variation of power, but for larger work.

Other techniques to compensate for oscillator drift or changes in the

loading conditions have been to use a feedback circuit [69U2] or to incorporate the transducer (piezoelectric) in the oscillator circuit itself.

These techniques have been incorporated in a few commercially available

bonding machines. The feedback circuit offers the opportunity of also being

used to monitor the making of a bond and thereby evaluate each bond as it is

made (see section 5.15).

Tl. Uthe, Uthe Technology, Inc., Sunnyvale California 94086, private communication.

42
4.6.4.3. Transducer and Coupler Assembly

The transducer converts the electrical oscillations from the power

supply to mechanical vibrational motion which is transmitted along and

amplified by a coupler assembly to whose end the bonding tool is fastened.

The coupler is pivoted about a nodal point of the transducer-coupler

assembly and the bonding force is applied through an adjustable dead weight

or spring.

If the mount for the assembly absorbs ultrasonic energy, power delivered to the bonding area is reduced and bonding is thereby affected.

Harman [70B2] encountered such energy absorption in testing several mounts.

By monitoring the tool tip vibration amplitude, different symptoms of such

absorption were observed: absorption at a subharmonic frequency; small,

irregular variations in the tool tip vibration amplitude; and a dependence

of the mechanical resonant frequency on applied ultrasonic power. Such

absorption was reduced by wrapping the mount screw threads with polytetrafluoroethylene plumbing type tape. This material is an excellent

damping material for ultrasonic vibrations. Johannesen [71J1] has

reported intermittent loss of ultrasonic energy in loose pivot bearings

in the mount.

transducer and coupler assembly

defective mount

4.6.4.4. Bonding Tool

The bonding tool is fastened near to the end and anti-node of the

coupler. The tool transmits the amplified vibration motion from the coupler

to the wire and bonding surface. The direction of the motion of the tool

tip is parallel to the bonding surface and the wire. Some early bonders

used a perpendicular vibrational mode [67R2] but this is no longer in use because of the difficulties encountered in counter balancing the coupler

assembly, transducer, and other attached accessories. Joshi [71J2] re- ported the use of vibration modes along each of three mutually perpendicular

directions, one being in the direction of the wire being bonded. He claimed

that the strongest bonds resulted when the tool vibration was in the direction of the wire.

The tool when energized with ultrasonic power from the coupler vibrates

not as a rigid beam but as a flexible beam with nodes and anti-nodes. The

relative vibrational amplitudes along tools of two principal lengths used are shown in figure 10. The tip of the tool oscillates thru a total distance of about 1 to 2 um for typical bonding conditions [70B2], [71B1].

Because the bonding tool is an integral part of the ultrasonic system, the

material, size, and position of the tool when installed have an effect on

tool vibration

vibration modes

vibrational

amplitude

*Bradfield [70B5], [70B6] has provided some background material on various types of transducers and their applications.

43
0.8

a> .co

c

0.6

g 0.4

LU

O

g 0.2

I i I J A 1 j £l_i i

0.4

o

c

86420 420

VIBRATION AMPLITUDE (relative values)

0.2 £

z

til o

o o

Figure 10. Typical vibrational amplitudes along a long and a short ultrasonic bonding tool

at tool extensions recommended by the manufacturer. Data was taken with a high resolution

capacitor microphone. The vibrational amplitude of the coupler (shaded) is also shown.

Amplitudes are in relative values. The tool tip amplitude for typical bonding conditions

is about from 1 to 2 urn. (From Harman [70B2], [70B3])

44
the operation of the system and hence the bonding operation. A given ultrasonic system operates best with tools of certain lengths installed with

certain extensions below the coupler (tool extension). The effect of a

tool extension

(effect of)

change in tool extension of as little as 15 mils (0.38 mm) on the relative

vibrational amplitude of the tool tip as measured by Harman [70B2] is shown

in figure 11. The figure also illustrates the vibrational amplitude enve- lope along the length of the tool.

Both the mechanical Q of the ultrasonic system and the relative tool

tip vibration amplitude may be affected by variations in the length of the

tool anchoring

bonding tool and its extension below the coupler, even for tools designed (effect of)

for the system [70B2], [70D2], Significant changes in the tool tip vibration amplitude have even been seen when the tool is replaced with the same

extension. This change was traced to the finish of the hole in the coupler,

through which the tool is fitted; the set-screw used to hold the tool in

place; and the torque used to tighten the set-screw. All appear to be able to affect the quality and position of the anchoring of the tool in the coupler

which in turn affects the vibration amplitude of the tool tip [70B2].

Tools are most often made of tungsten carbide. The mechanical charac- mechanical

characteristics

(WC tools) teristics of tungsten carbide depend on the series of processes used in its

fabrication. Accordingly, the performance of tungsten carbide tools in the

ultrasonic system is affected by variations in the fabrication processes.

These effects may be sufficiently significant to require changes in the

bonding schedule when tools made from a different batch are used. Therefore

it is prudent to keep separate tools made from different batches.

~, „.,,., . ,.,..,„ tool design

Tools are generally of wedge-type construction as shown in figure 12

with a wire-feed hole in the rear part of the tool through which the wire is figure 12 fed by a wire-clamp assembly (shown in figure 9). In some cases, a capillarytype tool is used to make stitch-type ultrasonic bonds and (when gold wire is

used) ultrasonic ball bonds. They are many designs for the wedge bonding

tools. Some of the design considerations are the shape of the bonding surface, radii of curvature of the front and back bonding surface edges, bonding surface finish, length of the bonding face of the tool (foot length),

and the size and location of the wire-feed hole. Optimization of tool

design is difficult because of the lack of adequate knowledge about the

effects of the various design features on the wire bond fabricated. Usually

a subjective decision is made, based on experience, tradition, and some cursory experimental work.

The bonding face is usually flat. Variations, primarily intended to tool face and

edges

minimize tool-to-wire slippage, include a groove perpendicular or parallel

to the direction of the wire. A problem with having a groove parallel to

*S. Bonis, Raytheon Corp., Sudbury, Mass. 01776, private communication.

45
.8

.7

.6

Q>

-C

o

c

X

o

2

UJ

O

I-

.2

\^

B

2 2

VIBRATION AMPLITUDE (relative values)

Figure 11. Vibrational amplitude along a long ultrasonic bonding tool for a tool extension

of 0.350 in (0.889 cm), At and a tool extension of 0.335 in (0.851 cm), B. The power

setting on the power supply was the same for both extensions but the frequency was slightly

different to achieve maximum amplitude in each case. Data were taken with a high-resolution

capacitor microphone. Rounding of the amplitude envelope at the tool ends are attributed

to edge effects. The vibrational amplitude of the end of the transducer coupler, which is

generally less than that of the tool, is shown in the center part of the figure. Amplitudes

are relative values. (From Harman [70B2])

46
^

/

TOE

FEED HOLE

HEEL

FOOT LENGTH

SECTION A -A

Figure 12. Sketches of the bonding end of an ultrasonic bonding tool.

the wire is the need to align the small diameter wire in the groove. To facilitate coupling between the tool and wire and to facilitate wire rupture

after the second bond has been made, the back or heel of the tool has a relatively sharp edge. However, this edge cannot be too sharp or excessive

deformation at the heel of the first bond can result.

The surface finish of the bonding face of the tool is also of importance in preventing undue slippage between the wire and tool. To avoid

inconsistent bonding characteristics due to such slippage, some have

recommended that new tools be "broken-in" by making a few hundred bonds

before using them on the production line. Plough et at. [69P1] have mentioned simply that the tool should have a rougher surface finish than the

wire. Uthe [69U1] has suggested that a sign that a good bond has been

made is that the part of the wire that had been in contact with the bonding tool should have a rough or granular texture, indicating that some bonding with the tool has occurred. Yet others claim that a smooth surface

is important, with no break-in period recommended; otherwise the wire tends

to stick to the tool and the bond is weakened when the tool is raised!

Use of tools with a short bonding face, or foot length, increases the

change of inadvertently applying too much pressure to the semiconductor

substrate and causing fracture. Also, Plough et al. [69P1] have reported

that the power setting is more critical when using tools with a short

surface finish

foot length

47
wire-feed hole

foot length and have recommended in agreement with Johannesen [71J1] the use

of a tool with as long a foot length as the bonding area will allow. For

1-mil diameter wire, foot lengths of at least 75 to 100 pm are generally

recommended

.

The size of the wire-feed hole and its location relative to the bonding

surface should be such that excessive rubbing of the wire against the tool is

avoided as the wire is paid out in the process of moving from the first bond

site to the second bond site [69P1]. Such rubbing tends to stress the heel

of the first bond or to nick and scrape the wire. This problem can be re- duced by moving the hole farther from the bonding face, but this introduces

wire alignment problems

.

4.6.4.5. Mechanical Stage for Transducer-Coupler Assembly

wire flexing

mechanical stage The mechanical stage for the transducer-coupler assembly is usually cam

operated to move the assembly and attached bonding tool in a vertical direction to either of two heights preselected to correspond to those of the die

and of the terminal bonding surfaces of a particular package. In some machines the stage moves the assembly backwards automatically in the direction of the second bond site after the first bond has been made to make a controlled wire loop.

The very nature of the bonding sequence after making the first bond subjects the wire at the heel of the first bond to considerable flexing which

can degrade the wire there [71G1] (see figure 7). Attempts have been made to moderate these stresses by moving the tool off the bond at an angle considerably less than 90°. Comments made privately indicate that these attempts

have not been entirely successful.

With some machines, the stage can also tip the wedge tool forward a few

degrees for the first bond to reduce the wire deformation at the heel of the

bond. The tool may also be tipped backwards on the second bond to increase

the wire deformation at the heel and thereby facilitate the break of the wire

there, after the bond has been made. A possible objection to the use of these

procedures is the increased sensitivity to minor misadjustments which because

of the smaller effective bonding surface of the tool may apply excessive

stress to the semiconductor substrate.

During fast operation the tool and associated assembly may not follow

the cams and as a result may vibrate or bounce. This can result in undesirable wire flexing. Such loss of control in tool movement is especially critical when the tool is lowered from the search height to the bonding surface.

Variations in tool impact velocity lead to variations in wire deformation,

before the ultrasonic power is turned on, and inconsistent bonding. For

tool tipping

operation speed

tool drop

The typical bonding forces used alone result in negligible wire deformation.

48
greater reliability Plough et al. [69P1], Davis [70D2], and Johannesen [71J1]

have recommended the operation of the bonding machine be slowed if such problems are encountered. Johannesen [71J1] has recommended a tool drop velocity

of 0.1 to 0.2 mils/ms (0.25 to 0.50 cm/s) . At greater velocities he noted

inconsistent bonding.

tool drop

4.6.4.6. Chuck and Movable Stage for Device

A chuck to hold the device package is part of a movable stage used for

positioning the bonding sites beneath the bonding tool. Johannesen [71J1]

has mentioned the difficulty in clamping the device package to attain the

required rigidity. The problem is most severe with packages made of brittle

materials because of the forces that must be applied.

Extraneous relative displacements between the tool and bonding surface

during bonding should be limited to significantly less than one-tenth the

wire diameter [71H1]. Such movements can be induced by a variety of sources

such as building vibrations, nearby machinery, and inadvertent bumping. It was reported that relative movements of 25 um or more could be induced in one bonding machine by only a slight movement of the hand while in contact with

the movable work stage holding the device. This stage was designed to be

rotated manually and the operator's hand would normally be in contact with

it during bonding [70B3]. In another experiment it was found that when lowfrequency vibrations of the bonding surface with an amplitude of about 25 um relative to the tool were induced during bonding, more than half the bonding

attempts were unsuccessful, either because of no bond adherence or because of

wire breakage [70B7].

chuck & movable

stage

package clamping

extraneous movement

4.6.4.7. Wire Clamps

The basic purpose of the wire clamp is to contact the wire and to feed

the wire forward under the wedge tool in preparation for making the first

bond and to pull back on the wire to break the wire at the heel of the second

bond.

Koshinz [67K1] mentioned the problem with the operation of some wire

clamps which because of their abrupt movements could cause mechanical transients in the wire. Laub and Mansour [69L2] describe a design which is intended to eliminate this problem. Leedy [70B3] reported that poorly finished

wire clamps could make indentations on the sides of the wire resulting in workhardening the wire and also embedment of contaminants in the wire at these

points. The latter is particularly important if the wire twists so that the

indented parts of the wire face the surface when the wire is bonded.

wire clamp

(use of)

problems

The text of [71J1] gives values of 1.0 to 2.0 x 10-lt mils/ms which is a typographical error (private communication with William Meyle, who presented the paper for Johannesen).

49
5. METHODS FOR TESTING AND EVALUATING WIRE BONDS

test methods

use of methods

test conditions

MIL-STD-883

discussion

analysis

correlation

:

5.1. Introduction

Methods used to test and evaluate wire bonds are described. To serve as

a quick reference for the reader, the methods are listed with brief, descrip

tive phrases in table 3. Some methods are used to test only the wire bond

while others also stress other components of the device and package as well

The latter fall in an environmental or environmental-simulation test category.

In general, the methods are intended to perform one or more of four functions:

(1) to optimize the process for making the wire bond, (2) to determine if this

process remains in control, (3) to obtain a measure of the quality of the wire

bond, and (4) to cull out those wire bonds that will fail under normal operational stress.

Where such information is available, the range of values for the test

conditions specified in standards and reported in the literature are tabulated with the description of the method. MIL-STD-883 [68D2] , [69D3], [6901]

has been influential to the extent that most of the test specifications

quoted in the literature are essentially identical to those of this military

standard. This is understandable considering that MIL-STD-883 was in part an outgrowth of test practices in the industry [69B8] and because of clear evolutionary links with the earlier (now revised) MIL-STD-750B [70D3] and MIL-STD202C [63D1], which in turn had widespread use. Additional influence has been

provided MIL-STD-883 by adoption of these test specifications by NASA as part

of its microcircuit line certification requirements [71N1] and in the purchase

of its microcircuits [69D3],

Also, each method is discussed in terms of its potential effectiveness

and practicality. These discussions include comments made in the literature

and general experience as related in interviews. Analyses of some of the

methods with regard to the stress that the test imposes on the wire bond have

been made and the results are used in the discussions of the relevant

methods.

Relatively little information has been found regarding the results of

testing the correlation between different test methods for the same type of

wire bond. What has been found is included in section 5.16.

50
Table 3 - Test Methods and Brief Descriptions

Air Blast (5.5) — A jet of gas is directed at the wire.

Bond Monitoring (5.15) — Some measure of the mechanical coupling

between the tool, wire, and metal film

is monitored during ultrasonic bonding.

Bond Interface Resistance (5 .14.2) — The contact or bond-interface resistance

is measured by a four-probe method.

Centrifuge (5.5) — A constant centrifugal force is applied

to the device.

i Electrical Continuity (5.14..J) — A small current through the bond is used

to test for electrical continuity.

Mechanical Shock (5.9) — A large, short-duration deceleration is

applied to the device.

Noise Measurement (5.14.4) — Noise is used as an indicator of contact

abnormalities

.

Pull (5.3) — Tensile forces produced by pulling the

wire by some means are used to stress the

wire bond either to some predetermined

tensile stress or to destruction.

Push (5.6) — A probe is used to apply a deflecting

force to the wire in a direction perpendicular to the plane of the wire and

bonds.

Short-Duration Stress (5.12) — The absorption of a short pulse of highenergy electrons is used to produce

thermally-induced tensile stress waves

in the device.

Shear (5.4) A probe, applied to the deformed wire

above the bond interface, is used to apply a shearing force at the bond

interface.

Temperature Cycling and Thermal Shock (5.13) —

Ultrasonic Stress (5.9)

Variable Frequency Vibration (5.10)

Vibration Fatigue (5.11)

Visual Inspection (5.2)

The device is exposed alternately to two temperature extremes.

Ultrasonic waves in a liquid medium are directed at the device.

The device is vibrated through some frequency range at a relatively constant, maximum acceleration.

The device is vibraJzed at a fixed frequency for a long period of time and at

a relatively small maximum acceleration.

The wire bond is examined to determine

if it conforms to predetermined criteria

of physical appearance and location.

51
visual

inspection

5.2. Visual Inspection

5.2.1. Description

widely used

test philosophies

table 4

reject criteria

bond placement

bonded wire

Visual inspection of the pre-encapsulated device is one of the most

widely used methods for evaluating wire bonds, as well as other parts of a device, such as its metallization. The examination of the wire bond is

usually performed with the use of a microscope under specified magnification

within the range of 30 X to 100 X and usually at specified angles of observation and lighting.

There have been two underlying philosophies in the selection of criteria

for visual inspection tests. One contains a greater reliance on (or expectation of) the test to detect the more subtle potential reliability problems as well as the gross defects. Consequently, the criteria of acceptance are more

detailed and more cosmetically orientated. The other embodies a greater

acceptance of the limitations of the test so the criteria are selected to cull out only the more obvious defects [66P1]. This approach is the more attractive from the point of view of testing efficiency and economics.

Four visual inspection tests including those of MIL-STD-883, are outlined in table 4 and represent varying responses to these points of view.

The higher the reliability required, the greater is the effort to use the

test to detect the more subtle potential problems. However, for the chance

of achieving greater reliability, the costs of reduced yield and additional

testing time and labor must be paid. Flexibility in terms of the reliability

sought is built into the visual inspection test of MIL-STD-883 [68D2]. Two levels or conditions A and B of the test are available to suit the reliability

need. NASA has adopted the modern version of the MID-STD-883 visual inspection test except that a greater magnification of 80 X is used [71N1].

The visual inspection tests outlined in table 4 are each divided into

three categories of reject criteria: bond placement, condition of the bond,

and condition of the wire. The comments in the paragraphs to follow related

to these criteria are meant to supplement the material in table 4. The bond placement specifications generally consider two potential

placement problems: One is concerned with the possibility of short-circuits

between adjacent bonds and between bonds and nearby metallizations. The

other is concerned with adhesion of bonds placed partially off the bonding

surface. Additional restrictions may be added to keep the bond on the die

from being located too close to the exit metallization stripe. The reason

for this restriction is that the bond can hide flaws or scratches such as produced by wafer probing which may have excessively reduced the conduction

path in the stripe area [6901],

The specifications for the condition of the bonded wire relate mainly

to the deformation of the wire at the bond and the bond contact area. The

52
rationale for allowing some reduction in the normal bond area is that the

strength of the bond between the wire and metallization is generally much

greater than the tensile strength of the most deformed part of the wire.

The purpose of the deformation specifications is primarily to cull out

bonds which have been over- or under-bonded. A reason a maximum length for

wedge or stitch bonds is specified is to cull out bonds where significant

relative parallel motion between the tool and bonding surface occurred

while the tool was in positive contact with the wire. The width of the bond

and any variation in it from bond to bond gives an indication of the degree

and uniformity of the bonding force. Variation in width of a given bond may

indicate misalignment of the tool relative to the bonding surface or wire.

Inspecting the height of the bonded wire is another way of determining

whether the deformation is within satisfactory limits and if any part of the

wire has been excessively deformed. To inspect the bond height, the device

must be viewed at an angle. wire condition

The specification for the condition of the wire relates to the location

of the wire relative to other parts of the device, the length of the wire

tail, reductions of the cross-sectional area or diameter of the wire, and

sharp bends in the wire. A maximum tail length is often specified for wedge

and stitch bonds. The tail length on the terminal is usually allowed to be

longer than the one on the die, where spatial tolerances are more critical.

Too long a tail is undesirable because of the possibility of its moving or even rupturing during a mechanical stress and thereby causing a short-circuit

Moreover, an excessively long tail would imply poor wire-feed control or a large elongation for the wire which, as discussed in section 4.2, would also

be undesirable. The criteria for excessive reduction of the cross-sectional

area in the wire adjacent to the bond or at any nicks, scratches or flaws are usually given in terms of a maximum allowed reduction in the wire diameter

but ocassionally they are given in terms of cross-section area which is more

difficult to judge. An indicator of possible overstress in the wire adjacent

to wedge and stitch bonds is a deviation, as viewed from above, of the wire

span from a straight line connecting the two bond sites. Thus a maximum permissible deviation is sometimes specified.

There are other methods,* besides using an optical microscope, to evaluate

wire bonds from an examination of their external appearance. The remaining SEM & *

vaJ methods paragraphs are devoted to the use of a scanning electron microscope (SEM) and

the use of x rays for this purpose.

*The recently introduced intensity spatial filtering technique [69W1] , [69W2] is not

applicable for wire bonds, however. Its use for detecting flaws in photomasks has been

detailed and its use for screening complex metallization patterns is indicated. However,

the relative lack of regularity in regard to the placement, shape, and routing of wire

bonds as well as the fact that they are not all in the same plane precludes the use of

this technique.

53
Table 4 - Visual Inspection Reject Criteria for Condition of Wire,

Condition of Bonded Wire, and Placement of Bond -

CONDITION

of

WIRE

REJECT CRITERIA

MIL-STD-883 Method 2010.1

Condition A Condition B 66L4 66P1 68R1

Magnification 30-50 30-50 50 20 -30

kd w

d <2D w

d <l0 wd <J3 wd ^0

w

d <L0 w

!— d p—

d <2D +

P

d <e t

P

d <3 mils

P d <e

-J2

d <e

P

d .<1 mil

P

h— •- —

H

D

T

die

L>2D

terminal

L>4D

die

L>2D

terminal

L>4D

L>2 mils L>2D

D <3/4D mD <3/4D m

D <3/4D m

D <3/4D m

D <1/2D

_m

D <2/3D m

TEAR' yes yes

R<2D

rmjnTTTn'w

REJECT CRITERIA

of

BONDED WIRE

MIL-STD-883 Method 2010.]

66L4 66P1 68R1 Condition A Condition B

Magnification 30-50 30-50 50 — -30

/<t

' \ffi

i ^^ W<2D; 6D<W W<2D; 6D<W W<2D; 5D<W

* r "iw W<1.2D; 3D<W

L<1.5D; 5D<W

W<1.2D; 3D<W

L<1.5D; 5D<W , L-L—J +

A, <1/2A. VA

i

Ab

I

'

: At

Rebonding on Die yes no yes sometimes

Rebonding .on Terminal no no yes

54
Table 4 - (Continued)

PLACEMENT

REJECT CRITERIA

of

BOND

MIL-STD-883 Method 2010.1

66L4 66P1 68R1

Condition A Condition B Magnification 30-50 30-50 50 20 -30

|~d>H d, <l/4 mil

b d <1/10 mil

b d <l/4 mil

b

(*- *m —|. d <l/4 mil m

d <1/10 mil m

d <l/4 mil m

If— df —-\ d

f <l/2 Wf+ d

f <l/10 mil+ d <2 mils 11 d

f <4/10 mil

\r**~\ d,<l mil

d

d <1/10 mil*

d h'.H d <.0 d <.0

<*w "^L-^i^t^^

d <0 d <0 w

d <0 w

d <0

1 w

I die11 V3/4A

b

terminal

A <A,_ m d

die

A <1/2A^ m d terminal

A <A^ m o

die

A <3/4A, ** m d A <1/2A++ m d

die

Am

<2/3A

b Am **

Ab A <1/2A§ § m b

Symbols

I 1 LZ=I

wire

metallization

TZ

Notes

vei"t.ical view

' wire at a distance greater

than 10 mils from its termination

§ at a magnification of 80 X

x if adjacent wedge bond to terminal + to farther edge of interconnecting metallization

stripe of width Wf

" also if bond is in interconnection metallization

stripe

n actually, if no separating

oxide is visible

a A < A. for hybrid circuits m b

* for wedge bonds

''"''for ball or stitch bonds

§§for ball bonds

d, -

d, -

other parts of device or package

non-contacting area

Notations

? edge of die

bond area bond-impression area

bond area over metallization

wire diameter. minimum thickness at deformation in wire

distance between boitds distance from bond to die

edge

distance from bond to farther edge of interconnection metallization

stripe

distance from bond to adjacent metallization

distance from wire to any part other than metallization or another wire

- distance of bond tail to adjacent metallization

- distance from wire to another wire or to metallization, as seen from a vertical view

- smallest distance at which wire will not touch other parts, if stressed

- length of bonded wire or of bond tail

- radius of curvature of bend in wire

- width or diameter of bonded wire

- width of metallization

stripe from

bonding area

55
Although the SEM has been used extensively in the failure analysis of

wire bonds its use, on a sampling basis, has only recently been introduced in

the NASA microcircuit line certification requirements [71N1]. The use of the

SEM in a screen test for wire bonds would be a powerful but a rather slow

tool to use. One problem with its use is the possible degradation of the

electrical characteristics of devices so examined [6801] . No degradation

effects are expected with use of a scanning electron mirror microscope [69C2],

However this technique is still in the development stage.

x ray X-ray examinations are used to detect defects in encapsulated devices

[68D2], With reference to wire bonds they are primarily used to view the

orientation of gold wire to see if the fabrication procedures subsequent to pre-encapsulated visual inspection or any previous test stresses have re- sulted in changes in the wire orientation which could constitute a potential

reliability problem. The use of an x-ray vidicon for this purpose has been

judged to be a highly valuable screening test for devices with gold wire

[66L4]. Because of the relative transparency of aluminum to x rays, however,

this technique is not useful for devices with aluminum wire bonds [66L4],

[68H4].

5.2.2. Discussion

discussion While visual inspection is widely used it is most effective when used in

conjunction with other tests. It cannot be expected to cull out all weak or

limitations defective wire bonds even if performed perfectly [6901], [69P1] , [69S4] . It

is used successfully to detect relatively gross visual defects but its success in detecting the more subtle defects is mixed [69S1] . In attempts

to do so it is also apt to be overly critical and reject acceptable wire

bonds, thereby needlessly reducing the yield.

criteria There are at least three basic limitations to the visual inspection test.

relevance

Firstly, the criteria in present visual inspection tests have not been fully

documented as to their relevance to the reliability of the wire bond. These

criteria are, essentially, an outgrowth of past experience, common sense, and

sometimes even guess. For example, in one instance when questioned for the reason for a criterion for an acceptable bond deformation that should be

allowed, the answer was that it was an educated guess.

Secondly, the test is subjective and suffers from human error. The more complex the product and the criteria, the less effective the inspector

[68H1]. It has been estimated that a visual inspection test performed by a

subjectivity

single operator is only about 80 percent effective. Thus, even if two non- correlative redundant tests were performed then only 96 percent of the possible defects would be detected. This is far from the kind of effectiveness

The implementation of such tests is no small problem because of psychological considerations,

56
required to detect one defective wire bond in 10,000 which for a 14-lead

device represents one device failure in about 700.

Thirdly, some defective wire bonds cannot be detected by visual means. The wire may be deformed as desired yet have essentially no adherence to the

bonding surface. This might occur, for example, because of surface contamination, poorly adhering metallization, or, for thermocompression bonds, too low a bonding temperature. The process of bonding may result in a fracture

of the semiconductor material under the bond yet the wire deformation at the

bond site may not appear unusual.

As a precautionary note it should be added that some of the criteria may

be expected to be dependent on the type of wire used. For example, the acceptable range of wire deformation may be dependent on the hardness of the wire

as well as its nominal diameter. Implicit in most visual inspection tests is

that 1 mil diameter wire is used. However, 1.5 mil and 0.7 mil diameter wires

are also in use. For 1.5 mil diameter wire, a wire deformation equal to or only slightly larger than the nominal wire diameter has been found to be

satisfactory while such a deformation would be inadequate for 1 mil diameter

wire. More work along the lines initiated by Kashiwabara et at. [69K1] is

needed to clarify the effects of these and other variables on the strength

of the wire bond and on an appropriate range of bond deformations.

detectability

precautionary note

5.3. Pull Test

5.3.1. Description

pull test

In its basic form, the pull test consists of pulling on the wire span

by some means (usually with a hook) with increasing force until either a predetermined force is applied or a rupture in the wire bond occurs. In the

latter case, the test is destructive; the pulling force required to produce

rupture is called the pull strength and it is used as a measure of quality

for the wire bond. In the former case, the test is meant to be nondegrading

as well as nondestructive to satisfactory wire bonds but destructive to the

rest. A variation of the destructive pull test is to cut the wire somewhere

along its span and then pull one of the wires at some angle with respect to the bonding surface and measure the force at which a rupture occurs. In some cases only one bond is made and tested. When only one bond is involved,

the test is called a single-bond pull test and the pull strength is equal to the tensile force in the wire. When two bonds are involved, it is called a double-bond pull test and the pull strength is generally not equal to the

tensile forces in the wire. Test specifications listed in table 5 summarize

the extent to which the conditions of the pull test are specified.

The pull test is used to (1) develop the bonding schedule by varying the

bonding machine parameters until acceptable pulling forces are attained,

pull strength

single-bond pull

double-bond pull

uses of method

57
data analysis

failure modes

(2) maintain control of the bonding process by periodically testing samples

of wire bonds that are produced, (3) determine the effect of other processes

or stresses on the wire bond by monitoring pulling forces of representative

samples, or (4) cull out weak wire bonds either by sampling a group which is

accepted or rejected on the basis of the performance of the sample, or by a 100 percent nondestructive test.

In the destructive pull test the pull strengths of a sample are taken to be representative of the group. The pull strength is recorded and, usually,

the location of the rupture as well. The pull strength is often expressed in

grams, although a grams-force unit is implied. To facilitate the analysis of

the data, the distribution of pull strengths can be displayed in a histogram.

The magnitude of the pulling force at the peak in the distribution displayed

in the histogram gives an indication of the general ruggedness of the wire

bond while the spread of the distribution indicates the uniformity of a group, how well the fabrication processes are under control, and how uniform

are the constituent materials.

The failure mode, or location of rupture, reveals the weakest point in

the wire bond for the kind of stress imposed. The most common point of rupture is in the wire adjacent to one of the bonds, because of the reduced

cross-sectional area or mechanical strength there. Thus, the pull test is

Table 5 - Pull Test Specifications and Conditions

Source Gauge Accuracy Rate of Pull Probe Material Diameter of

Wire Probe

71N2*

71B4

70A1

±0.5 gf for 0-10 gf

±1,0 gf for 0-100 gf

±0.05 gf (load cell)

"slowly"

1.0-77 gf/s

steel, tungsten

molybdenum

5 mil

(0.13 mm)

5 mil

(0.13 mm)

69K1

69B5

2 mm/min

2 mil

(0.05 mm)

68D2 +0, -0.5 gf for 0-10 gf

+0, -1.0 gf for 10-50 gf

67S3

67H1

66R1

0.77 mm/min

"slowly"

0.2-0.5 in/min

(0.13-0.5 mm/min)

Test is performed under a microscope at a magnification of 45 X.

58
more a test of the deformation of the wire than a test of bond adherance, as has been pointed out by Howell and Kanz [65H2], and others. If the rupture

occurs at the point of contact of the wire with the probe (usually a hook)

then the probe itself may be damaging the wire and invalidating the test. This is most apt to occur after high temperature stress which has annealed

the wire.

The great majority of pull tests use a hook fabricated from a wire, pulling element

typically tungsten, to pull on the wire span. The other end of the wire is

attached to a gauge to measure the applied force. The wire diameter at the

hook is typically between 40 to 80 urn (1.5 to 3.0 mil) and a primary concern

is that the hook not damage the wire during the test. In cases where it is

felt that the height of the wire span is insufficient to position a suitable

hook, an

M

' L" shaped v probe or a pointed, straight-wire probe has been used.

• f ,» a v rate of pull

The rate of pull is a parameter that is occasionally specified for the

pull test. It is given either in terms of the speed of the pulling element

[66R1], [67S3], [69K1], or the rate of increase of the force, as measured at the pulling element [70B8], In some pull tests, the only specification is

that the force must be applied "slowly" [67H1], [71N2].

Test apparatus for making both destructive and nondestructive pull testing machines

tests is commercially available. A design for incorporating a nondestructive

pull test in a bonding machine making special "eyelet" thermocompression

bonds was reported by Baker and Jones [66B4] . At the time of the report, a prototype wire-feed in which the wire is retracted by rotating the wire

spool had been constructed for 2-mil diameter aluminum wire. Nothing since

has been seen in the literature about any further development or use of this

design. Present-day bonding machines do not feature it. sinple-bond pull

The purpose of a single-bond pull test is usually to better control the

angle of pull (with respect to the bond and contact surface) and also thereby

minimize flexing and hence weakening the deformed wire adjacent to the bond. In the case of a completed wire bond, the wire is severed somewhere along

its span and one half of the wire bond is tested. However, a means of

gripping the wire is required in this test. Wasson [65W1] has described gripping wire

the use of special tweezers to grip the wire. Adams and Anderson [68A1]

have recommended the use of a black wax. A separate heating element is

used to liquify the wax at the end of a probe which engulfs the wire by

capillary action. They report that the tensile strength of the wax is

sufficient to measure pull strengths up to 150 grams-force. More recently,

Harman [69B5] has described the use of high-tensile-strength hot-melt glues,

with discrete melting-points, at the tip of a nichrome wire-loop probe.

The glue is melted by passing current through the wire loop which has been

electrolytically thinned so that most of the joule heating will occur there.

The tensile strength and adherence of these glues are sufficient to test

59
grip stress

purpose

1-mil diameter aluminum wire if about 0.13 mm length of the wire is bonded

with the glue. The glues do not bond as well to gold wire so that a longer

length of gold wire must be immersed in the glue.

The use of wax or glue minimizes the grip-stress on the wire. Such

stress is greater for a mechanical grip (by tweezers for example) . Yet,

while the grip-stress may be larger, Dudderar [69D1] indicated that the rapid

strain hardening minimizes its effect and thereby serves to avoid wire rupture near the region of mechanical contact.

analysis 5.3.2. Analysis

This analysis is intended to provide relationships between the pull

strengths and the actual force in the wire which can result in rupture and

thereby also indicate the sensitivity of the test to the parameters of the

test. By resolving the forces in the pull test as shown in figure 13, the

resolution of ,.,_,- _ ,, ,. . , - , .,

forces applxed pull force, F, is related to the wire tensile force on the terminal

side, F . by ' wt' '

cos(e, - <f>) F = F x (5.3.1) wt sin(e

d + e t

)

and to the wire tensile force on the die side, F ,, by

cos(e^ + (J)) = V x wd

F 3 = F x £ . (5.3.2)

sin(6

t + e d

)

The angles 6 and 6, are the angles that the wire makes with the bonding surt d faces of the terminal and die, respectively. The pulling probe is inclined

at an angle (j> with respect to a normal to the bonding surfaces. The case where the direction of pull is at some angle from the plane of the loop is

not included in the analysis.

8 dependence The wire tensile force on the terminal side of the hook is greater than

the applied force at the hook if 2e ^ + 6 - <f> <90 deg. The ratio F /F as

d t wt

a function of 6 is given in figure 14 for various fixed ratios of 6,/6 and

figure 14 t & a ^ t for <f> = 0. The ratio F ,/F as a function of 0, may be obtain by interchanging the subscripts d and t. These curves indicate how dramatically the

tensile force in the wire may differ from the force applied by the hook.

Wire bonds with different contact angles may have entirely different tensile

forces in the wire for the same pulling force applied at the hook.

Similar equations have appeared earlier [70B7], Equivalent expressions for <|> = have been

derived by Donald J. Manus, Texas Instruments, private communication.

60
Figure 13. Geometric variables for the double-bond pull test.

I

Figure 14. Dependence of Fwt /F on 9 t for various ratios 6^ to 6 t

. The curves are for the

case <j) - 0. Horizontal bars above and below the curves for 26^ = 9 t and 0<i = 20£ show the

effect of changing <j> to plus and minus 5 deg, respectively. To obtain the dependence of

Fwt/F, interchange everywhere the subscripts d and t and change the sign of <J>.

61
wt wd

The dependence of F /F on the magnitude of cj> is also indicated in fig

<j> dependence Wt ure 14. The effect of varying <j> by ±5 deg is shown by the horizontal bars o either side of the curves for G./G = 1/2 and 2. Positive values of 6 reel t suit in a larger F . The dependence of F on 4 increases with the ratio wt r wt

sin(8, - c}>)/sin(6^ + 6,). The symmetry between F and F , is maintained if

d t d

J wt wd

the sign of <j> is reversed with the interchange of the subscripts d and t. The ratio of the wire tensile force at the terminal side to that at the

die side of the probe is given by

F cos (0 - cf>) -^ = * (5.3.3)

Fwd cos(6

t + 40

Because the terminal bonding surface is usually higher than the die, 9 is

usually less than 6,, and in this case, the tensile force in the wire is

greater on the die side of the probe than on the terminal side.

h, H, o, d Although the geometrical dependence is more easily visualized in terms

dependence

of the contact angles, these angles are difficult to measure in practice.

The contact angles depend on the height, h, of the wire span above the

terminal contact surface, the height difference, H, between the die and

terminal contact surfaces, and the horizontal distances ad and (1 - a)d, from

the bonds to the point at which the wire span is contacted by the pulling

probe. These quantities are shown in figure 13. Expressing equations

(5.3.1) and (5.3.2) in terms of these quantities we have

\1 + ( —J (1 - a) cos <j> + (

J sin <j> f -ill—^—

L

—\-jU. JL,

wt 1 aH

1 + —

h

l~ (1 - a) 2 d

2

/ H\ r 1 (h + H) 2 \ h/ L

h

cos <S> sin

F =f{- < h + H> i ^J= d Jl (5i3%5)

wd 1 aH

1 + —

h

For a pulling force normal to the bonding surfaces (<j> = 0) , these equations

simplify to:

(1 -

1 +

a) / /ad\2

»«t * '

i —ST

^ "V*i * (-) + \-l

h

62
F , = F wd aH I

1 + —

h

(1 - a) 2 d

2

(h + H) 2 (5.3.7)

For a pulling force applied at mid-span and normal to a single-level bonding

surface (<j> = 0, H = 0, and a = 1/2) these equations simplify further to:

? =F=F=- V1 + W/2h) 2. w wt wd „ » (5.3.8)

Work is underway to confirm experimentally the dependence of pull strength on the geometry of the wire bond which is given by the above equations

[71B1], [71B4], [72B1], [72B2].

Figures 15 and 16 have been provided to assist in the practical use of

the relationships given in equations (5.3.6) thru (5.3.8). In figure 15 the

ratios F/F and F /F as functions of d/h are graphed for a single-level

(H = 0) , double-bond pull test where the pulling probe is located at midspan. The ratios F /F and F ,/F as functions of d/h for different values

wt wd

of H/h are graphed in figures 16a and 16b, respectively. Note that here

the probe is located at a distance of one-fourth the bond separation from

the bond on the terminal which is a rough estimate of where the apex of

the wire loop would be and where the probe would normally be located in a two-level, double-bond pull test.

To use the pull test in evaluating the control of the bonding process

or the effect of factors which alter the strength of the wire bond, it is

desirable to know the sensitivity of the pull strength to changes in the test variables, assuming that the wire tensile force necessary to cause rupture is constant and the failure mode does not change. With this information one may determine the extent to which the distribution of pull

strength observed is actually due to variations in the quality of the wire

bonds being tested. In selecting a particular configuration for the wire

bond, it may be possible to select a configuration in which the sensitivity

of the pull strength to those variables which can be least controlled in

making a series of bonds is minimized.

The measure used here to describe this sensitivity is the normalized

variation in the pull strength, AF, where it is assumed that the tensile

force in the wire to rupture the wire bond is unaltered by variations in

the geometrical variables of the wire bond and the test.

Fi - F

experimental

confirmation

figures 15, 16

F/F ; F /F w w

sensitivity to variance

AF

AF = (5.3.9)

where F represents the pull strength obtained when each of these variables

has a selected value and F} represents the pull strength obtained when the

63
Figure 15. F/Fw

and Fw/F as functions of d/h are graphed, for a single-level, double-bond

pull test.

64
Figure 16a. Fwt /F as a function of d/h for different values of H/h and for a = 1/4 is

graphed, for a two-level, double-bond pull test.

5 —

H/h«0.

J I I I I I I I L J L

10 14 d/h

18 22 26

Figure 16b. Fw<j/F as a function of d/h for different values of H/h and for a. = 1/4 is

graphed, for a two-level, double-bond pull test.

65
AF(t)

values of one or more of these variables have been altered. The magnitude

of AF will depend on which side of the pulling probe rupture occurs. To distinguish between cases where rupture occurs on the terminal and where it occurs on the die side of the pulling probe the symbols AF(t) and AF(d),

respectively, are used.

The following expressions for AF(t) and AF(d) in terms of h, H, a, were

derived from equations (5.3.4) and (5.3.5):

I

AF(t) =

1 +

hi

h + H

(1 - a) cos <J> H sin <(> d

aH

1 + —

h

hi + Hi

(1 - CL [) COS lf< ^

+" " - Sill (J)

x , 1 ai J ./Mi*

(5.3.10)

axH], H h

1 H 1 H a cos <j> sin

h x h d

AF(d ) AF(d) = x x aH H: h l 1 H 1 + — ai cos (\>i - — sin <$>i h h x d x c

(1 - a) 2 d

2

(h + H) 2

(1 - a x ) 2 d

2

(hi + Hi) :

- 1, (5.3.11)

where the subscript, 1, denotes the altered magnitude of the variable and

subscript denotes the initial value.

dependence of Implicit in the above expressions is the dependence of h on a for a

h on a given wire bond. To assist in visualizing how h will vary with a, it is useful to consider that the part of the wire in contact with the probe lies on an ellipse whose foci are located at the bonds.

figures 17-22 The curves shown in figures 17 thru 22 display the sensitivity of AT to changes in the parameters a, h, d, H, and <J>. In each case only the quantity

in question was varied.

66
-10°

CURVE (-«)« h+H

1 1 2 2

3 4

AF(t)

20%

- - -10%

-•-20%

Figure 17. Dependence of AF(t) on pulling angle, 4>i » ^ or different values of

(1 - a)d/(h + H) and for (j> m 0. The

dependence of AF(d) may be obtained by

reflecting curves about the vertical

axis and changing the running parameter

to ad/h.

CURVE CtH/h

1 »l

2 4

3 1 4 0.5

AF(t), AF(d)

20%

0.8

CURVE

od_ (i-q)d

h ' h+H

1 »l

2 2

3 1

"^-20% 1.20%

Figure 18. Dependence of AF(t) on the

ratio H^/H for different values of aH/h

and for <J> = 0. The variation of AF(d)

is more complex but it will be about

equal to or less than that shown for

AF(t).

Figure 19. Dependence of AF(t) and

AF(d) on di/d for <p = . The running

parameter for AF(t) and AF(d) are ad/h and (1 - a)d/(h + H)

,

respectively.

67
AF(t),AF(d)

Figure 20. Dependence of AF(t) and

AF(d) on h^/h for <j> = 0. The running

parameters for AF(t) are aH/h and

ad/h; for AF(d) they are aH/h and

(1 - a)d/(h + H).

CURVE d/h H/d

1 2 2 10 3 2 0.4 4 10 0.4

CURVE ad (l-a)d

h ' h + H aH

h

1 »l

2 » 1 1 3 2 4 2 1 S 1

CURVE d/h H/d

1 2

2 10 3 2 0.4 4 10 0.4

AF(t)

207o

10%

Figure 21. Dependence of AF(t) on a^/a

for a = 0.5, <j> = 0, and different

values of d/h and H/d. The dependence

of h on a is included. The dependence

of AF(d) will be similar, particularly

as the squares of (1 - a)d/(h + H) and

ad/h become large with respect to unity.

Figure 22. Dependence of AF(t) on a]_/a

fora=0.25, <j) = 0, and different

values of d/h and H/d. The dependence

of h on a is included. The dependence

of AF(d) will be similar, particularly

as the squares of (1 - a)d/(h + H) and

ad/h become large with respect to unity.

68
5.3.3. Discussion discussion

The pull test is perhaps the most widely used test for wire bonds. The

wide use test is fast, easy to perform, requires little equipment, and it provides a number, the pull strength. The ease with which this number can be obtained

may be a reason for the lack of attention that is often given to details of

the test that may have a significant effect on the result.

Despite its long use and its importance in rating the quality of wire incomplete

specification

bonds, no satisfactory standard pull test method is available. When the use

of a pull test is indicated, the test conditions are under-specified. When

the results of a pull test are reported, not only are the test conditions

incompletely specified but so also are the pertinent physical parameters of

the wire bond. Thus, in the typical situation some average pull strength

is quoted and histograms may be shown, but because of the lack of other information the data are of limited usefulness and may even be misleading. When

such data are used to qualify parts or to compare different wire bonds, con- siderable confusion and misrepresentation can occur and the technically more

sophisticated competitor, as well as the user, may be penalized.

specification To maximize the usefulness of pull test data especially when used to needs evaluate and compare wire bonds it is important to specify fully the wire

bonds tested and the test conditions as well as the failure mode. The

following several paragraphs discuss many of the items that need to be

specified for the wire bond and the pull -test conditions.

To adequately identify the wire bond tested, such information as the

wire material, its mechanical characteristics and cross-sectional dimensions,

the bond type, the metal film(s) and the substrates at both bonding sites are needed because all may affect the pull strength measured. A particularly

important wire characteristic that has a direct effect on the pull strength

is the tensile strength. More useful is the tensile strength of the wire at the time of the test. Exposure of the wire to elevated temperature anneals

the wire and reduces its tensile strength significantly. For example,

Pankratz and Collins [70P1] reported that the tensile strength of aluminum

wire with 1 percent silicon added was reduced to about 60 percent and to about 40 percent of its initial value after an exposure of one hour to elevated temperatures of 200°C and 250°C, respectively. Fully annealed wire

can have a tensile strength of as little as 20 percent of its initial value.

It is necessary to have information about the average shape of the wire

span.* Changes in the shape can significantly affect the pull strength measured. To indicate the magnitude of the dependence of the pull strength on

the shape of the wire span in a double bond -pull test, consider two wire

tensile strength

wire span

*In cases where the elongation of the wire is large, such as in annealed wire, it is necessary to take this into account when providing information about the shape of the wire span.

69
bond efficiency

span variance

gauge accuracy

bonds for which the same tensile force in the wire is required to produce

rupture (so that if a single-bond pull test were performed both would indicate the same pull strength) . Let the contact angles that the wire makes

with each bonding surface be 10 degrees for one wire bond and 30 degrees for

the other. The pull strength measured for the one with contact angles of

30 degrees will be about three times as large as for the other (see figure

14).*

In order to evaluate and compare the pull strengths of different wire

bonds, some means of normalizing the data is needed. In such an attempt,

Riben et al. [67R1] used the term bond efficiency or the ratio of the tensile

force in the wire at rupture to the tensile strength of the wire.''' Riben

et at. reported bond efficiencies of from 0.32 to 0.63 while more recently

Cox et al. [70C1] reported bond efficiencies of from 0.8 to 0.9, all for

aluminum wire ultrasonic bonds. These data were for single-bond tests where

the pull strength is equal to the tensile force in the wire at rupture. For

a double-bond pull test, information about the geometry of the wire bond and

the location and direction of pull is needed to relate the pull strength measured to the tensile force in the wire.§ This kind of information is rarely

if ever provided (or even recorded) with the consequent inability to compare

different wire bonds reported in the literature in terms of their pull

strengths.

The ranges of the geometrical parameters used to describe the shape of

the wire span are also important. Such information can be used to estimate

the contribution that variations in shape may have had on the width of the

pull strength distribution. (This may be done with the help of figures 17-

22).

Virtually the only specification that is given for the pull test

apparatus, as may be seen in table 5, is in regard to the accuracy of the

gram-force gauge reading [68D2], [71N2], But, the requirement for accuracy

in the gram-force gauge is virtually meaningless considering the incompleteness of the specification of the wire bond. For example, MIL-STD-883 requires that the accuracy be +0, -0.5 grams-force in the range of to 10

grams-force. Assuming the same tensile force along the wire is required to produce rupture, variations considerably larger than 0.5 grams-force are to be expected for the different wire bond geometries that may be used in

For wire bonds made on the same plane, 10 degrees and 30 degrees correspond to values for

the ratio of loop height to bond separation (h/d) of about 0.1 and 0.3, respectively.

"fjoshi [71J2] used the term bond quality factor which he defined as the tensile force in the

bonded wire at rupture, when the wire is pulled parallel and close to the bonding surface,

divided by the initial wire tensile strength.

§ The effect on the measured pull strength of geometrical factors such as the wire span, loop

height, relative height difference of the two bonds, and the location of the pulling probe

on the wire span has been presented in detail in the preceeding section.

70
microelectronic devices. Another possible source of error will be present

if the gauge, which typically uses a lever arm to transmit the pulling

force, is not rotated during the test to maintain the pulling force perpendicular to the lever arm. There has been some concern that the speed of the test could affect rate or p

the magnitude of the pull strength measured. Riben et at. [66R1] mentioned observing a decrease in single-bond pull strength (magnitude not

given) by increasing the speed from 0.2 in/min to 0.5 in/min (^ 5.1 to ^ 12.7 mm/min) . However, Leedy and Main [71B4] reported no dependence of

pull strength or pull rates in the range of 1 to 77 gf/s for wire bonds with

1-mil diameter aluminum wire bonded ultrasonically to an aluminum film on silicon. The range of loading rate reported by Leedy and Main is roughly

equivalent to a pulling rate range of 0.39 to 30 mm/min which includes the

range of pulling rates in which Riben et at. reported a dependence. The

higher rate of 30 mm/min may be comparable with the speed of some pull test machines used in the industry where the pull may be likened to a jerk. It should be added that the wire bonds used by Leedy and Main were constructed

so that rupture occurred in the heel of the bond. The independence of the

pull strength on pull speed in the range reported may not hold for wire

bonds where the failure mode is rupture or peel at the bond interface or where the two bonds are on different levels [71B4]

.

To better define measurement conditions in the pull test in any pull rate specification

future experiments and to quantify test specification with regard to pull rate, the alternate ways of defining pull rate recommended in the

ASTM standard method for tension testing [70A6] should be considered.

Five ways are listed: (1) rate of straining, (2) rate of stressing,

(3) elapsed time for completing part or all of the test, (4) speed of

the moving head during the test, and (5) speed of the moving head when

not under load. All but the last measure would be useful as a means

of quantifying the pull rate. Sufficient information should be provided so that given one way it would be possible to determine the rate

by another. Relations are provided in Appendix B which related the

first four of the above measures for the pull rate. Some elements in the design of the pulling probe, should be men- probe design

tioned. The shape and diameter of the probe at the pulling end are

important with regard to possible damage to the wire. The stiffness

and the suspension of the probe are important to the control of the

direction of the pulling force and the control of the location of the

probe on the wire.

The effect on the pull strength of variations in the pulling angle out of plane

in the vertical plane containing the two bonds is indicated in figure 17.

If the direction of pull is outside this plane such a twisting stress

71
probe location

nondestructive

test

evaluation

should have a degrading effect on the measured pull strength. Leedy and Main

in preliminary attempts to measure such degradation for single-level, aluminun

ultrasonic wire bonds found only small if any degradation in pull strength for

angles as large as 20 degrees out of the plane of the wire loop.

The ability to locate precisely the pulling probe on the wire span is

also important.. The orientation of the wire may make it difficult to position the hook at the desired location and then pull the wire perpendicularly

to the bonding surfaces without having the hook slide along the wire. This can occur when the bonding surfaces are at different levels. It may be possible to circumvent this difficulty by pulling in a direction which is perpendicular to the plane of the two bonds, but it may be inconvenient to manipulate the device so that the proper angle of pull is obtained. Slipping of

the hook may also occur when testing ultrasonic wire bonds which have a slight bend in the wire span as a result of the fabrication procedure (see

figure 7) . The location of this bend may not correspond to the point where

it is desired to position the probe. The procedure of deforming the loop

initially in an attempt to create a bend at the desired location to prevent

the probe from slipping is tedious and not always successful. Furthermore,

this involves additional flexing of the wire.

Thus far the comments made in this part have been related to destructive

pull tests. Many of the comments, however, also apply to the nondestructive

pull test which has been mentioned as a 100 percent screen test by Slemmons

[69S1] and Ang et al. [69A1], To substantiate the claim that the test is

non-degrading to those that pass, the authors show that after a group of

wire bonds has been so tested and then pulled to destruction, the bond

strength frequency distribution is merely truncated at the preselected stress level.

Polcari and Bowe [71P3] have reported the results of some preliminary

evaluations of a nondestructive pull test. The test vehicles were 1-mil

diameter aluminum, ultrasonic wire bonds in uncapped 14-lead integrated circuit devices made by one manufacturer. They found that prestressing once to

2 gf produced no significant statistical differences in the mean and in the

standard deviation of the pull strengths of wire bonds stored for 250 hours

at temperatures of 150, 200, and 250°C. However, multiple prestressing to

2 gf did appear to be degrading to wire bonds. They concluded by emphasizing

that while the nondestructive test could be a valuable reliability tool it

must be recognized that the proper use and adjustment of the tester would be

of critical importance. Furthermore, if the nondestructive pull test were

to be implemented, they would recommend that control over the variation in

K. 0. Leedy and C. A. Main, National Bureau of Standards, Washington, D. C. 20234, to be

published.

72
loop heights be exercised by the manufacturer to minimize variations in

the stress imposed by the test. Although there seems to be considerable interest in a nondestructive

test there is also skepticism that the test is actually non-degrading. The

idea of using devices whose wire spans have been pulled and, in the process,

altered in shape is new and disturbing to many. Yet, there does not seem

to be any reluctance about using devices after having been stressed in a centrifuge test. Another matter about which uncertainty exists is the level

of stress required to cull out the weak wire bonds without degrading the

remaining ones. Another reservation raised about the nondestructive pull test is the

time required to make the test. In general the test will take approximately

the same amount of time that it takes to make a wire bond. If in fact the

nondestructive test can be used effectively, the extra testing time would

be very well spent, especially for high reliability products.

reservations

5.4. Shear Test

5.4.1. Description

shear test

The shear test involves the use of a probe to apply a horizontal force

to the wire bond. The shearing force necessary to rupture the bond is taken

as a measure of the quality of bond adherence. This test was developed primarily to measure bond adhesion of gold ball bonds to metallization films.

Other types of bonds do not generally offer a large enough area to probe.

The probe described in the literature [67A1], [67G1] is a pin with a blunt end. The use of a hypodermic needle tip, ground flat at the end,

has also been observed during a laboratory visit. The exposed capillary

at the tip accommodates the ball bond in a way which deforms it less than

a pin-type probe and thereby a more uniform loading is applied to the bond.

MIL-STD-883 [68D2] specifies only that the probe be of such shape that it causes a minimum of cutting at the probe edge when applied to the bond.

Shear strengths as large as 100 grams-force have been reported [67G1].

Typical ranges reported extend from about 10 to 70 grams-force [67A1],

[67G1], [69L1].

5.4.2. Discussion

adherence test

probe

shear strengths

discussion

Judging from the literature and personal contacts, the shear test is

not widely used. There appear to be several problems with the method. One

is the fundamental difficulty of designing a probe small enough to contact

the deformed wire above the bond and yet strong enough to apply sufficient

shearing force. Another is the operator dependence of the test resulting

in part from the difficulty of contacting the bonded wire reproducibility.

little used

problems

73
metallization

test

air blast test

During one laboratory visit it was reported that a variation of from 20 to 40

grams- force could be expected in the shearing force when different operators

measured equivalent wire bonds.

Nevertheless the method has been used to monitor and optimize the adhesion between the bond, metallization layer (s) and the substrate as well as determine the effect of various stresses on this adhesion [67A1]. Gill and

Workman [67G1] found the shear test to be the only effective method for predicting poor adherence between films of gold and molybdenum. In particular

they found that adherence problems are indicated by abnormally low shear

strengths or when the gold film is sheared off with the wire to expose the

molybdenum film.

5.5. Air Blast Test

5.5.1. Description

An air jet from a nozzle whose diameter is much larger than the wire

bond tested is directed at the device in a way which will apply a generally

lateral thrust to the wire loop. The force of the air jet is adjusted so that poorly adhering bonds or wire badly deformed at the heel will be ruptured without degrading wire bonds of satisfactory strength.

The test is most suitable for use when the wire bonds are arranged

symmetrically on the device. Either the device is rotated under the jet of

air, directed at some angle from the vertical, or the device is fixed and the

nozzle is moved about the device. Electrical tests follow to detect those

wire bonds which have been ruptured.

Design information has generally been considered to be proprietary. No

reports or published papers have been found which describe the apparatus or the method

.

5.5.2. Discussion

reservations No evidence, pro or con, is available to evaluate the effectiveness of

the method. An often expressed reservation about the method is the difficulty of quantifying the stress imposed on the wire bond. Variables that

affect this stress are the orientation of the wire bonds and the semiconductor

die, and the design of the package. The force exerted by the air jet in turn

is affected by such characteristics as the nozzle design, air pressure, angle

of incidence, and the distance between the nozzle and the wire bonds. These

effects have not been studied. In the face of the consequent uncertainties,

the tendency is to minimize the stress level of the test and thereby reduce

its potential effectiveness. Plough et al. [69P1] found that while the method

would detect very weak wire bonds, the air jet could also cause metalfatiguing oscillations in satisfactory wire bonds.

74
push test

new method

5.6. Push Test

5.6.1. Description

The new and potentially nondegrading method, a push test, has been

suggested by Floyd. The test involves the application of a force to the

apex of the wire span in a direction perpendicular to the wire and parallel

with the bonding surfaces. Initial, exploratory use of the method has been on ultrasonic wire bonds. An appropriately selected force is applied by a rejection

wire probe. The magnitude of the force is indicated by the deflection of cr terion

the probe.'" The criterion for rejecting the wire bond is either a large

displacement of the wire loop at its apex (indicating plastic flow of an excessively deformed wire adjacent the bond), rupture of the wire at the

heel (indicating gross deformation there), or rupture of the bond (indicating poor adherence)

.

5.6.2. Discussion

The push test is similar to the air blast test in that a lateral force . blast vs to the wire loop is used to cull out weak wire bonds. The push test offers Push

greater control over the applied stress and greater flexibility in that it can

be used conveniently even when the wire bonds are not symmetrically arranged.

The testing rate can be comparable with that of the nondestructive pull test,

but the test has the potential advantage of not altering the shape of the

wire span and so allowing less chance for degradation. On the other hand

there is then no evidence that the wire bonds have been tested. Considering

the potential advantages of the method further exploration is warranted.

A basic problem is that the displacement of the wire due to the appli- , , cation of the force at near mid-span may not be a sufficiently sensitive measure of the condition of the heel of the bond which is usually the weakest part of the wire bond. Except for very weak wire bonds the deflection

should occur mainly in the wire between the heel and the point of application,

because of its much greater length. It remains to be seen if the push test can be made to be sufficiently sensitive.

There are other problems which must be recognized. In its present form,

the applied force is controlled by deflecting a wire probe a specified amount.

Both this deflection and the maximum deflection of the wire span, of the order

of a wire diameter or less, must be estimated by the operator during the test.

*H. L. Floyd, Jr., Sandia Corp., Albuquerque, N. M. , private communication. He related that

the idea initially came from an unidentified engineer who used a whisker from his beard as

a probe and called it the "whisker test".

"^Forces as large as tenths of grams-force have been used. H. L. Floyd, Jr., private

communication.

75
ultrasonic

stress test

This may be a significant source for variability in the use of the method.

The dependence of the stress to the wire bond on the shape and length of the

wire span is present in this as in many other tests in use. Thus the maximum

allowed deflection of the wire span may have to be adjusted to take the geometry of the wire bond into consideration.

5.7. Ultrasonic Stress Test

focused US energy

needs evaluation

cleaning bath

centrifuge test

Knollman et at. [69K4] have described tests on capped devices using equipment in which a controlled level of focused ultrasonic energy is directed at

the under-side of the device package. Two modes of stress were used: a pulsed mode and a cw mode with power density of up to 500 W/cm2 for a duration

of 20 to 200 us, and up to 32 W/cm2 for 1 to 10 s, respectively. The most

prominent failure for the pulsed mode was rupture at the bond interface while

for the cw mode it was rupture at the wire adjacent to the bonds.

Not enough work has been done to evaluate the possibilities of this test

method. So far the exploration of the method has been entirely empirical. A

more theoretical approach is needed to provide a better understanding of the

nature of the stress on the wire bond and to determine how it is affected by

such things as configuration of the wire bond and the design of the package.

A much less desirable method of stressing wire bonds ultrasonically is

to place the uncapped device in a commercially available ultrasonic cleaning

bath. An immediate problem with this method is the large variation of ultrasonic power density with location in the bath.

5.8. Centrifuge Test

5.8.1. Description

purpose

conditions

specified

The centrifuge test subjects a device to a constant acceleration in

order to determine the ability of a device component to sustain the centrifugal force imposed. When the intent is to test wire bonds, the device is

oriented so that the centrifugal stress either (1) ruptures poorly adhering

bonds or (2) displaces wires which have an excessive loop or which are improperly placed so that they will make electrical contact with adjacent parts of

the device. This rupture or displacement is detected by electrical tests following the centrifuge test. The pertinent conditions specified for the test are the maximum acceleration (expressed in gravity units, g) , the duration of the test at maximum

acceleration, the orientation of the device with respect to the direction of

the centrifugal force, any prior stresses, and the measurements to be made

after the test to detect failure. The typical range of centrifugal accelerations used to test 1-mil diameter gold wire bonds is from 30,000 to 50,000 g's.

A typical duration at the maximum stress level is 1 min in any one orientation.

76
The orientations chosen are those in which the centrifugal force is directed either away from, toward, or parallel with the bonded surfaces.

The test is often preceded by other tests intended to weaken unreliable

wire bonds and thereby promote their failure in the centrifuge test. Test conditions given in standard methods or reported in the literature

are listed in table 6.

Table 6 - Conditions Used or Recommended in Centrifuge Test Methods

Source Acceleration (1,000 g's) Directions of Centrifugal

Force*

Duration of

Maximum

Stress (min)

70D3 + not specified six 1§

69B2 20 , 40, 60, 80, 100 *1 1 6901 30

69S4 10-100 Yj or six

68B1 10-30; 1 200

68D2+ 20* six or less 1 68I1#

5, 10, 20, 50, 100* six or less 11/6

68R2 20 six

67G1 20, 100 x l» X2 , Zi, Z 2 , Yi

66G1 30 Y l» Y2» z i, Z2

66L4 20 Yi, Y2 1§

66P1 20 Yi, Y2 65C5 20-40

65R1 20 six

65S2 40, 70, 100

64U1 10, 20 Y!Y2 , X-Z plane

*Along X, Y, Z coordinate axes of device as specified in MIL-STD-883, for example;

Y axis is meant to be normal to the bonding surface, subscript 1 indicates direction away from surface, subscript 2 indicates opposite direction. If all six

directions are given, the word "six" is entered.

+MIL-STD-750B.

^Duration between zero and maximum acceleration must be at least 20 seconds.

XA11 six directions for reliability testing; Yj direction for screen test,

^MIL-STD-883.

AT Shall apply unless otherwise specified.

3,000; 5,000; 7,500; 10,000; 12,500 g's.

*IEC Recommended Method.

aLower acceleration conditions included.

Other values listed are 500; 1,000; 1,500;

77
tensile force

catenary span

contacts

5.8.2. Analysis

The calculated tensile stress on the wire bond resulting from a centricalculation fugal force directed away from the bonded surfaces is provided in this section. These calculations allow a better evaluation of the usefulness of a centrifuge test in testing wire bonds of a particular design than the calculations reported by O'Conner [6901, figure 3] which were for a simpler geometry and are applicable only for small values of the ratio of wire loop

height to bond separation, h/d.

It is assumed that the centrifugal force is sufficiently large so that

the wire span takes the shape of a catenary. The parameters of the system

are illustrated in figure 13 (section 5.3.3).

tensile force at l- t "^y De shown that the tensile forces in the wire, F and F , , at the

contact points to the terminal and die respectively are in grams-force:

F fc = p Trr 2 G (a + h) (5.8.1) wt

F , = p Trr 2 G (a + h + H) (5.8.2) wa

where a = for d % 2 (H + h),

4h (1 + /l + (H/h)) + 2H

and where p = density of the wire (g/cm 3

)

,

r = radius of the wire (cm)

,

h = vertical distance between the terminal contact surface and the

peak of the wire loop (cm)

,

H = vertical distance between the terminal and die contact surfaces

(cm),

d = horizontal distance between bonds (cm), and

G = centrifugal acceleration (in units of gravity)

.

approximations Using the approximate value for a will introduce an error of less than

10 percent in F and F even for d/(h + H) as small as 2, which is an unusual case. An exact value for a can be obtained from the relation

h + H + a = a cosh (D/2a) , where D/2 is the lateral distance between the

bond at the die and the apex of the wire loop. Actually, the greatest un- certainty may result from the estimate of h, which is the distance between

the apex of the wire loop and the terminal contact surface after the centrifugal forces have deformed the loop to describe the catenary curve. A useful graphical representation of equation (5.8.1) (using exact

values for a) is shown in figure 23. Here, the tensile force in the wire

adjacent to the bonds of a single-level, 1-mil diameter, gold wire bond subjected to a centrifugal force of 30,000 g's is shown as a function of d for

figure 23

78
Figure 23. Tensile force, Fw , in the wire adjacent to the bonds in a single-level, 1-mil

diameter, gold wire bond due to a centrifugal force of 30,000 g's directed perpendicularly

away from the bonding surface is graphed as a function of d for different values of d/h. Fw for a given value of d is a minimum when d/h ^ 3. Values for Fw for accelerations other

than 30,000 g's may be obtained by multiplying the value for Fw by the ratio of the acceleration of interest to 30,000 g's. Values for Fw for 1-mil diameter aluminum wire bonds may

be obtained by multiplying Fw by 0.14.

79
example different values of d/h. The figure shows, for example, that for a wire bond

with a bond separation of 0.15 cm (^ 60 mils) and a loop height of 0.015 cm

(^ 6.0 mils) the tensile force in the gold wire will be about 0.55 gf . Such

a tensile stress could be produced in a pull test if the hook, placed at mid

span, were pulled with a force of 0.25 gf.

explanation of The way in which the tensile stress on the wire depends on the variables

w

varia lon of the wire bond may be seen by first considering the single level case wher

H = and where d 1 2h so that the approximate value for a may be used. The

tensile forces at the wire ends are equal and given by

\8h /

rd 2 Fw

2( p Trr 2 G ( — + h) . (5.8.3)

s

For a constant acceleration, F decreases as h is increased from a value small w

in comparison with d. This follows because the angle that the wire makes with

the horizontal, 6, increases and thereby decreases the component of the cen- trifugal force along the wire. When h becomes comparable with d, F increases w

with h. This follows because the increase in wire length for a unit increase

in h is larger and the consequent increase in mass and hence centrifugal

force will dominate. The dependence of the tension in the wire at the bonds

as H is increased from zero follows in a similar way from the way changes in

d, h, and H affect the angles 6 and 6, and the length of the wire.

5.8.3. Discussion

widely used The centrifuge test is widely used in the industry. It is regarded by

many as a reliable, nondestructive method of screening for defective gold

wire bonds, especially when preceded by other stress tests such as temperature cycling, mechanical shock, or vibration tests. 0' Conner [6901] has

stress levels indicated that while stress levels due to accelerations of from 40,000 g's to 60,000 g's are more effective, a 30,000 g's acceleration level is specified

in MIL-STD-883 because such equipment is generally available and the test at

this level can be performed without special jigging. If accelerations much

greater than 50,000 g's are used, there is a risk of damaging other components

of the device.

It is generally agreed that the test is not useful for screening alumi- num wire bonds because of the much lower density of aluminum. However, there

imitations are those in the industry who believe that even for gold wire the centrifuge

test is acceptable for weeding out only grossly defective wire bonds [67A1].

Basic to their objections is that the test does not impose a significant

stress unless excessively large accelerations are used.

80
5.9. Mechanical Shock Test mechanical

shock test

5.9.1. Description

The purpose of the mechanical shock method is to evaluate the structural

integrity of components and to determine their suitability for use where they

may be subjected to non-repetitive mechanical shocks such as in handling or in operation.

In a typical test the device is first accelerated either by free-fall or by pneumatic means and then brought to a sudden halt on striking an impact

pad. The test conditions that may be specified are the maximum deceleration

(in gravity units); the duration of the impact or shock pulse-width; the

direction of the stress (usually along one or more of the principal axes of

the device package) ; and the number of shocks per direction. MIL-STD-750B

[70D3] is one of the rare domestic tests which specifies the shape of the

shock pulse. The IEC [6711] specifies the use of one of three basic pulse

shapes: saw-tooth, half-sine, and trapezoidal.

Test conditions of standard methods and methods used and reported in the

literature are listed in table 7.

5.9.2. Discussion

As will be discussed later, the stress imposed by a mechanical shock test

is less than twice the stress imposed by a centrifuge test at a constant

acceleration equal to the peak shock-induced deceleration and directed in the same direction, neglecting the uncertain effect of the number of vibrations

that are generated. Because the largest peak acceleration listed in table 7

is 6,000 g's and considering the small stress that a centrifuge test at twice that acceleration would impose, the mechanical shock test will not

stress most wire bonds greatly.

To make the test more severe by increasing the amplitude of the shock

can result in deformation of the package at impact unless special supporting

structures are used. The fact that some package designs do not readily allow means for support along one or more axes increases the inconvenience of the

method. Shurtleff [69S4], however, expressed satisfaction with the method

for culling out weak bonds. He suggested that electrical monitoring during

the shock would improve the effectiveness of the test but the increased

difficulty and expense incurred would normally preclude the use of any

monitoring.

Lombardi et al. [66L4] suggested that the shock test represents a com- promise between the centrifuge and variable-frequency vibration tests in

that it does not usually impose on the wire as large a force as the centrifuge

purpose

test conditions

stress level

low stress

stress limitation

shock vs. centri- fuge & variable

frequency

81
Table 7 - Conditions Used or Recommended in Mechanical Shock Test Methods

Source Peak

Acceleration

(1,000 g's)

Shock Pulse

Width (ms)

Shock Pulse

Shape

Number

of Shocks

(each direction)

Direction

of

Shock

70D3 §§ § + half -sine X X 69B2 1.0 5 to 60 *1 +

69SA 1.5 to 6.0 5 to 10 * 68D2XX 1.5« 0.1 to 1.0 1 5 * 67I1++ 3.0, 1.5, 1.0,

0.5,...#

0.2, 0.5, 1.0,

1.0,..."

saw tooth,

half-sine, and

trapezoidal**

3 tt

* Along both directions of the three major coordinate axes of the device as specified

in, for example, MIL-STD-883 [68D2], for a total of six directions.

*r Y axis is meant to be normal to the bonding surface; subscript 1 indicates the direction away from the surface.

§ Value to be specified, ±20 percent.

* To be specified. + Value to be specified. Tolerances: for x i 2 is, the greater of ± 0.6 ms or ±15 percent of t; for t £ 2 ms, the greater of ±0.1 ms or ± 30 percent of t.

If Shall apply unless otherwise specified. Other values listed are 500; 3,000; 5,000;

10,000; 20,000; and 30,000 g's.

v Other values are listed. Tolerances: for saw-tooth and half -sine shaped pulse, ±15

percent; for trapezoidal shaped pulse, ±20 percent.

" Other values are listed.

**Tolerance bands are provided.

"'"'"Each direction of three mutually perpendicular axes chosen to impose maximum stress

(total of six directions).

§§MIL-STD-750B.

xxMIL-STD-883.

^IEC Recommended Method. ________——————

test nor as wide a frequency spectrum as the variable-frequency test. He

reported further that several manufacturers believed that one would not cull

out potential device failures with the shock test which the centrifuge or variable-frequency tests had not removed. Thus, the use of the mechanical

shock test would be redundant if the other two tests were used. However, if

testing is to be minimized at the sacrifice of screening effectiveness, the

shock test could be used in their stead. The validity of these propositions

is not supported by any experimental data.

Actually for a half-sine excitation pulse with a 0.1 ms pulse width, as an example, the

shock test would generate significant energy at frequencies higher than the 2 kHz maximum

used in the variable-frequency test methods.

82
Considering the test in terms of the mechanical response of the wire

bond, the use of the term shock is somewhat of a misnomer. The mechanical

response time of the wire span is either comparable with or less than the

shortest mechanical pulse duration used (0.1 ms) . (For an estimate of the

lowest resonant frequency of a wire bond see Appendix A.) In essence, the

basic stresses to the system result from the amplitude and the number of

vibrations induced in the wire span by the mechanical pulse. The number

of the vibrations depends on the damping of the wire loop. Thus, if the

wire has been annealed, through earlier exposure to high temperatures,

fewer oscillations with smaller amplitudes occur and the resulting flexure

stress may be less. However, irreversible changes occur in annealed wire

at smaller bending and torsional stresses than unannealed wire.

To discuss the test in terms of the magnitude of the induced deflection

and stress, it is useful to define a dynamic-to-static deflection or stress

ratio, K:

"shock"

K = maximum wire deflection or stress by a mechanical pulse

wire deflection or stress induced if peak acceleration

of mechanical pulse were applied statically

As a rule of thumb, the upper bound for K is 2 [48F1], [65R3, p. 368], [6711].

How much less than 2 it is depends on the shape and duration of the pulse and on the lowest resonant frequency of the wire loop. The typical case occurs when the lowest resonant frequency of the wire loop is greater than the in- verse of the duration of the excitation pulse (compare table 7 and figure A2

of Appendix A). In this case, K tends to be larger (but is bounded by 2) the

lower the resonant frequency. The test is therefore most severe for mechanical shocks in a direction normal to the plane of the wire loop, because the

lowest resonant frequency in this case is that for the lateral vibration mode

perpendicular to the plane of the wire span.i"

Changes in the shape of the excitation pulse will produce significant

differences in K. Consequently the IEC in its recommended test method has

required the specification of the shape of the shock pulse. For the purpose

of reproducibility, the use of a saw-tooth shaped pulse is suggested because

it produces the least variation in vibration amplitude with changes in pulse

duration and resonant frequency [6711].

types of stress

deflection

stress

shock-pulse

shape

*If the lowest resonant frequency is less than the inverse of the duration of the excitation

pulse, as may be the case for gold wire bonds with large bond separations, K decreases to zero the lower this resonant frequency is.

i"The resonant frequency for the lateral vibration mode is significantly less than those of

the other two modes, except for wire loops where h/d is small. For the straight wire case

(h/d = 0) the resonant frequencies for the lateral and extensional modes are equal, as may

been seen in figure Al.

83
ripples

variable, frequency vibra- tion test

purpose

test conditions

The IEC method [6711] also stresses the need to minimize any ripples on

the shock pulse in order to maximize the reproducibility of the method; ripple

is considered to be the most significant distortion of the shock-pulse shape.

It is therefore with some concern that such ripples are noted in a display of

a typical impact shock pulse from a pneumatic shock tester described recently

in the literature [71K1].

5.10. Variable Frequency Vibration Test

5.10.1. Description

The purpose of the variable frequency vibration test is to excite and

rupture any device components, including wire bonds, having a resonant frequency within the range swept. The selection of the frequency range is subject to the constraints of test equipment available and an estimate of the frequency components of the kinds of shock and vibrations the device may encounter

in its life. In a typical test the device is vibrated sinusoidally through a range of frequencies for one or more cycles. A minimum cycle duration is

usually specified. The test is repeated along each of the three principal

axes of the device. The maximum acceleration (in gravity units) is also

specified, sometimes with a limitation on the maximum amplitude at the lowest

vibration frequencies thus reducing the peak acceleration there. The test

conditions in standard methods and methods reported in the literature are

listed in table 8.

5.10.2. Discussion

test ineffective The variable frequency vibration test does not impose a significant

stress on most wire bonds and so the test is generally ineffective. The

lowest resonant frequency of most wire bonds is many times larger than the

maximum vibration frequencies listed in table 8 (see Appendix A) . Hence, the

dynamic deflection of the wire and the induced stress will at most be only

slightly larger than that induced if a constant acceleration equal to the peak

acceleration were applied [65R3, pp. 368 and 370]. The maximum acceleration

used is usually less than 100 g's which would produce vibrations in the wire

with a negligible amplitude. Krieg and Murfin [69K2] introduced a scale factor to transform a system to a larger size where intuition can perhaps be of

greater help in estimating the severity of the stress imposed. For example,

using their concept of a scale factor, a 1-mil diameter aluminum wire, 0.127 cm

long and subjected to 100 g's peak acceleration at a driving frequency of 2 kHz

is equivalent to a 2.5-cm diameter aluminum rod, 125 cm long being driven at a frequency of 2 Hz to a peak acceleration of 0.1 g's (equivalent to a maximum

displacement of about 1.2 cm), hardly a stressful condition.

84
Table 8 - Conditions Used or Recommended in Variable Frequency Vibration Test

Source Peak Frequency Sweep Rate Duration or Number Direction

Acceleration Range (Hz) of Cycles in of

(g's) (logarithmic scan) Each Direction Vibration

70D3* - 20

69S4 2-20

6802* 20, 50, 70^

6611** 1, 2, 5, 10, 15,

20, 30, 50*

66L4 30

100-2000

20-2000

20-2000

10-5000

10-2000

10-500+ *"

10-2000

^L 4 min/cycle

il 4 min/cycle 4

"» one octave/min 1/2, 1 1/2, 6,

30, 150 hr

4 min/cycle

Along the three major coordinate axes of the device package as specified in, for example,

MIL-STD-883.

t Three mutually perpendicular axes chosen to impose maximum stress.

* Below about 80, 127, and 150 Hz for the respective peak accelerations, the total excur- sion is limited to 0.06 inches.

x Below 60 Hz the total excursions are limited to those at 60 Hz for the peak acceleration

levels listed. Vibrational amplitude tolerances are given for "control" and "fixing"

points. + More and narrower ranges are included.

' A linear sweep may be used if the sweep rate does not exceed one octave per minute.

# MIL-STD-750B.

MIL-STD-883

.

IEC Recommended Method

.

Shurtleff [69S4] indicated this basic problem of low stress. MacKenzie

and Carpenter [67M1] reported that while vibration tests at peak accelerations of up to 70 g's caused no failures in gold wire bonds to aluminum films,

they did cause a significant number of package failures.

5.11. Vibration Fatigue Test

5.11.1. Description

The purpose of the vibration fatigue test is to determine if the com- ponents of the device, including wire bonds, can withstand relatively long

periods of low-frequency sinusoidal oscillations without metal fatigue.

The test involves vibrating the device at a low, fixed frequency, for many

hours, along each of the three principal axes of the device package. The

maximum acceleration (in gravity units) or vibrational amplitude is specified. The test conditions in standard methods and methods reported in

the literature are listed in table 9.

vibration

fatigue test

purpose

test conditions

85
Table 9 - Conditions Used or Recommended in Vibration Fatigue Tests

Source Peak Frequency Duration Direction of

Acceleration (Hz) Per Direction Vibration

(g's) (hr)

70D3X L 20 60 ± 20 1 32 ± 8 * 69S4 2-10 60 100+

68D2+ 20, 50, 70 § 60 ± 20 1 32 * 8 * 65C5 30 60 32 —

Along the three major coordinate axes of the device packages as specified

in, for example, MIL-STD-883.

^"Whether duration is for each direction or for all directions is not specified.

c Acceptable alternative to 20 g's is a constant total excursion of 0.06

inches

.

XMIL-STD-750B.

+MIL-STD-883 .

.

5.11.2. Discussion

The vibration fatigue test is an ineffective way to cull out weak wire

bonds for the reasons stated under the variable frequency vibration test. Shurtleff [69S4] adds that the duration of the test and equipment maintenance during the test makes it prohibitively expensive as a 100 percent

screen test.

short-duration

stress pulse

test 5.12. Short-Duration Stress Pulse Tests

The use of short-duration stress pulses to apply large tensile stresses

to the wire bond has been explored by Floyd [68F1]. The stress pulses are

, ..". generated by the absorption of pulses of high energy electrons in an absorbing

plate fastened to the base of the device package. The deposition occurs in a time short compared with the time that a stress inhomogeneity will decay; thus

it may be assumed to occur instantaneously and to generate a thermally induced

compressive stress pulse. Because the momentum of the incident electrons is

negligible, a tensile stress wave follows as the compressive stress wave

propagates so as to maintain zero total momentum. Upon reaching a free surface, such as the top surface of a wire bond or a semiconductor chip, the

compressive stress pulse is reflected as a tensile wave. This results in an amplified tensile condition in regions where the reflected tensile wave overspallation laps the incident tensile wave. In these regions the tensile stress may be

high enough to exceed the spall threshold along some plane and cause device

86
failure. Sufficiently great tensile stresses have been generated with adequate resolution to cause a range of failures in a number of devices, from

rupture of the wire bonds at the terminal and the semiconductor die to fracturing of the semiconductor die itself. These results were duplicated

by the use of a stress-wave-analyzing computer code [68F1], This method

has been referred to elsewhere [69P1] as an electronic anvil test. Another method for generating short stress pulses is by the use of

explosively accelerated plates, sometimes called the plate-slap or the

flyer-plate method. Floyd [68F1] indicated that the momentum transfer in

such techniques can cause spurious stresses which make analysis difficult.

5.13. Temperature Cycling and Thermal Shock

5.13.1. Description

anvil test

plate-slap,

flyer-plate

methods

temperature cycling and thermal shock test

Exposing the device alternately to two temperature extremes is a com- monly used method to test the ability of the wire bonds as well as other

parts of the device to withstand the mechanical stresses that result from

differences in the thermal coefficients of expansion of the constituent

parts. The main justification for such tests is the expectation that the

device will be exposed to similar temperature extremes prior to and during

its operational life.

whether the test is referred to as thermal shock or temperature cycling

depends on the transfer time between the two temperature extremes. The

transfer time for shock tests is of the order of seconds while for cycling

tests it is usually of the order of minutes. Air chambers are usually used

in temperature cycling tests while liquid baths are used in thermal shock

tests. For temperature cycling tests, the minimum time spent at each temperature extreme is longer in order to allow the device under test to attain

the ambient temperature.

The extremes of temperature used usually fall within the range of -196

to +300°C. A common low-temperature extreme is about -65°C while a greater

variation in the high-temperature extreme is seen: gold-wire, thermocompression bonds to aluminum films are usually tested to 125°C or 150°C, aluminum

wire bonds to about 200°C, and gold-wire to gold-film bonds have been tested

to 300°C. Some temperature cycling tests also include an intermediate

period near room temperature. The minimum number of cycles used in a screening test is about 5 to 15. When tests are used to stress wire bonds

to failure, hundreds or more cycles may be used. The test specifications

of MIL-STD-883 and MIL-STD-750B for the temperature cycling and thermal

shock tests are listed in table 10. The conditions of most of the other

tests used and described in the literature are similar to one of these.

purpose

transfer time

test conditions

87
Table 10 - Conditions Used or Recommended in Temperature Cycling and Thermal Shock Tests

Source High : Low Maximum Minimum Time Minimum Number

Temperature Extremes Transition at of

(°C) Time Temperature Extreme Cycles

68D2§ 125« - 55+0* 5 min 10 min 10

150+5 -65+0 5 min 10 min 10

200+J

" 65+0 5 min 10 min 10

350+5 - 65+°

— c5 5 min 10 min 10

500+J

-65+0 5 min 10 min 10

68D2* 100+5

-0 o+o

-5 10 s 5 min 15

TM+l - 55+0

-5 10 s 5 min 15

150+5 - 65+0 10 s 5 min 15

200+5 " 65-°5 10 s 5 min 15

150+5 -195+5 10 s 5 min 15

200+5 -0

-195+5 -5

10 s 5 min 15

70D3+ 100+0

— b °^o

5 3 s 15 s; 5 s

+

5

ioo+o 0+5

-0 10 s 5 min 5

Cycles include a period of not more than 5 minutes at an intermediate temperature of

25° +i°: C.

4-At the high and low temperature extreme, respectively.

§MIL-STD-883 , method 1010.

xMIL-STD-883 , method 1011.

+MIL-STD-750B. method 1056.1. conditions A and B. ,

ordering of tests

monitored

thermal cycle

power cycling

The ordering of a series of tests is important. If temperature cycling

and thermal shock tests are used as part of a series of screening tests then

it is important that they be performed before mechanical stress tests (such

as the centrifuge and vibration tests) are made so that the wire bonds weakened by the temperature stress tests can be culled out [66G1], [66L4], [6901]

The use of a "monitored thermal cycle" test [71H2] is discussed in

section 5.14.3. In this test the device is also exposed to temperature

extremes. The thermal-induced stresses are used to dislodge or shift poorly

adhering wire bonds which are detected by performing electrical continuity

tests.Power cycling tests produce stresses in wire bonds similar to those induced during temperature cycling and thermal shock tests. However, the

88
stresses as a result of wire flexure are less in a power cycling test for

the same maximum device temperature. This follows because the expansion

of only part of the wire, nearest the semiconductor chip where the power

is dissipated, contributes appreciably to the wire flexure.

5.13.2. Analysis

The tensile and flexing stresses resulting from the differential thermal expansion of the wire bond and the device header will be discussed in

this section. These stresses are important considerations in not only temperature cycling and thermal shock tests, but also in power cycling tests where there is sufficient time during the power-on and power-off parts of

the cycle for the system to approach thermal equilibrium.

bonding and To estimate the order of magnitude of stress levels that may be gene- tensile stress rated in the wire adjacent to the bonds, Krieg and Murfin [69K2] made approximate calculations of the maximum bending and membrane stresses in the

wire (away from the bonds) of specific wire-span configurations. These

calculations did not include the wire ends because of computational difficulties. The results of these calculations showed that these stresses in- crease as the contact angle is decreased. Their expressions do not extrapolate to the case of a straight wire. In this case, if there is no buckling,

the tensile stress is given'*' by a = E(3 - 3 ) (T-T), where E is the modu- w s o lus of elasticity, 3 and 3 are the thermal coefficients of expansion of the

wire and surface, respectively, and T and T are the initial and final tem- peratures, respectively. For a 1 mil-diameter aluminum wire where E £ 7 x 108 gf/cm2 (70 GN/m2

) and 3 - 3 = 15 x 10-6 °C_1

, the tensile force will

increase by about 0.05 grams-force per degree Celsius decrease in

temperature.

An important stress, especially as the number of cycles is increased, is

the flexing that the wire undergoes adjacent the bonds. The most obvious flexing stress

manifestation of this flexing is the change in loop height, h. Gaffney et at. [68G2] observed loop height changes of as much as several wire diameters in

small-signal transistors during low-repetition-rate pulsed power operation.

They concluded that the wire-fracture failures they found were due to flexinginduced fatigue of the wire. Recently, Ravi and Philofsky [72R1] documented

*An analysis has been made by Phillips [72B1] , [72B3] of the thermal- expansion induced

Changes in loop height of a 500-mW, 50-mA silicon transistor in a TO-18 package which are

produced by the dissipation of power in the semiconductor die as would occur in a power

cycling test. His analysis has shown that the temperature gradient is linear along the

wire between the die and the terminal in an unencapsulated device, indicating that the

radiation and convection heat losses from the wire are negligible. Joule heating in the

wire was assumed to be negligible at the relatively low current levels considered.

'''For example, see P. R. Lancaster and D. Mitchell, The Mechanics of Mateviats, McGraw-Hill,

New York, 1967, p. 40.

89
change in h

the formation and propagation of microcracks and eventual failure by fracture

at the heel of ultrasonic bonds in a dynamic SEM display of wire flexing as a result of such power-on-off operation.

Gaffney et at. [68G2] also calculated the change in the loop height for

a circular-shaped wire loop as a result of the thermal expansion of the wire,

uniformly heated, for a fixed horizontal bond separation. Their expression

is rearranged below

V

2n(3n + 1)

1 + 3 (3n +5) (n - 1)

W

(T - T ) (5.13.1)

where n is the ratio of the wire length at T = T to the bond separation.

This expression is developed from the relationship 3s - 8/hz + (d/2) z - d

which relates the circular-arc length of the wire, s, to the loop height and

bond separation to within 2 percent for ratios of h/d as large as 0.5. Using

this relationship to obtain

/3

h = — d/(3n +5) (n - 1) , (5.13.2)

°

8

n may be obtained in terms of h and d which when substituted into equation

(5.13.1) provides an expression in terms of quantities that may be easily

measured

h i°fTK9 a h„"Vl + - 17+ ) 3w

(T - T

q

) (5.13.3)

For most practical cases and where d/h is not excessively large, the square

root term may be expanded to obtain

h = h

o

1 / 3 d

2 \ - (7 + ) 3 ( 8 \ 2 h

2 /

w 1+- J 7 + -) 3„(T - Tq

) (5.13.4)

In these cases, h is directly proportional to T - T and the percentage change

in h increases as h decreases. Similarly, it may also be shown by differentiating equation (5.13.1) that dh/dT decreases with increasing values of

T - T .

* The results of these calculations were presented at the conference but not subsequently

published

.

90
: While the change in loop height offers a convenient measure of flexing,

I the more meaningful measure is the change in the contact angles. The re- change in contact angle

mainder of this section will deal with the results of calculations made of

the thermal-induced changes in contact angle of a single-level wire bond

i assuming the configuration of the wire space remains unchanged under expansion or contraction as a result of changes in the ambient temperature.

The changes in the contact angles were calculated for two cases: where the wire span forms the two equal arms of an isosceles triangle and

where it forms a circular arc. These two shapes were chosen in an attempt

to bracket a range of loop shapes that might be encountered. It was found

that for a circular arc shaped loop

sin ib sin \p _ o^

* *o

[1 - (6w

- 3s

) (T - T

Q )], (5.13.5)

and for an angular shaped loop,

cos i|> = cos lp [1 - (6 - 3 ) (T - T )], (5.13.6)

o w s o

where iJj, \b = the final and initial contact angles, respectively,

\b =2 tan-1 (2h/d) , for the circular arc span,

\b = tan-1 (2h/d), for the angular shaped span,

h = loop height (cm)

,

d = bond separation (cm)

,

3,3 = the thermal coefficient of expansion of the wire and bonding

surface meterial, respectively (°C -1

), and

T, T = the final and initial ambient temperature, respectively (°C)

.

The calculations incorporated the fact that both 3 and 3 are very much w s smaller than unity.

Using the above equations the change in contact angle, \p - \b , versus

h /d is graphed in figure 24 for the two wire-span shapes, for a differential expansion between the bonding wire and bonding surface, (3 - 3 )

it (T - T ), equal to 0.004. Scales are also, provided in figure 24 to show

the percentage change of h /d for both wire-span shapes. The thermalo o

j induced change in contact angle and h/d increases in magnitude as the initial contact angles or as h /d decreases. For a given initial contact

o o angle or h /d , the change in contact angle is less for the angular span

*For a wire bond with aluminum wire (3 - 24 * 10~6 °C-1) and a base material with a 3g of

4 x 10-6 °C _1

, the value of 0.004 for

W

(3 - 3 ) (T - T ) represents a temperature change of

200°C. w s o

91
-^ 1° ' Angular span

2s, >

CIRCULAR ARC SPAN

(0w-£s >(T-To)= 0.004

oJO^^

" <o2

o

z

<

a:&

o

-p

o

UJf

<

z

o

*? -|I50

-t300

4100 J

50

25

- 10

5 J

2.5

100

50

25

2.5

1.5

0.05 0.10 0.15 0.20 0.25

h /d

0.30

Figure 24. Change in contact angle, i|i - i|/ , versus initial value of the ratio of the wireloop height to the wire span, h /d , for a differential expansion between the wire and the

bonding surface, (^ - B s ) (T - T )

, of .004. Scales parallel to the if> - ^ axis are provided to show the percentage change in h /d for each wire-span shape.

92
than for the circular-arc span. The difference between the curves for the

two wire-span shapes shows how sensitive is the temperature-induced change

in contact angle and h/d on the shape of the wire span. To assist in estimating the thermal-induced change in contact angle,

\\> - i\> (in degrees), from an observed change in loop height, the following

expression for the circular-arc span is given for conditions when

(h /d ) « 1 and d SdQ. Then,

* - <l» * h

* - *

o £ 4 180 h h - h ^ , o h - h

- 230 (5.13.7)

for a circular arc span. The expression for the angular arc span is the same except the numerical factor should be halved. The above expression

shows that for the conditions stated, the thermal-induced change in contact

angles is proportional to the thermal-induced change in loop height.

Phillips [71B5] reported earlier that changes in contact angle were

approximately proportional to the change in loop height for wire flexing

resulting from low-repetition-rate power on-off cycling conditions.

5.13.3. Discussion

Contradictory evaluations have been made of temperature cycling and

thermal shock tests. Bell [69B1] considered these tests as effective for

screening bond failures as did Hakim and Reich [65H3] who also regarded

them as effective in comparing production processes or products. On the

other hand, Plough et at. [69P1] dismissed the usefulness of thermal

shock tests on ultrasonic wire bonds, saying that they did not induce

failures even in wire bonds known to be weak. On one thing there is

general agreement. The thermal shock test is regarded as the more severe

test and is therefore sometimes used to accelerate failure of those devices

which would eventually fail as a result of temperature cycling tests. However, there has been no experimental work published to document this

opinion.

The reason for disagreements about the usefulness of temperature

cycling and thermal shock tests is that all wire bonds are not equally

stressed by these tests. As discussed in the analysis section, an important stress especially as the number of cycles is increased is the flexing

that the wire undergoes adjacent to the bonds. Obviously the magnitude of

the flexing is affected by the differences in the thermal coefficients of

expansion of the constituent parts of the wire bond and by the shape of the

wire span. The magnitude of this flexing increases for smaller values of

h /d or smaller initial contact angles for a given change in temperature

discussion

contradictory

evaluation

wire bonds not stressed equally

dependence on

h /d

93
(see figure 24). Phillips [72B1], [72B2] reported that a group of 95 transistors fabricated with aluminum wire bonds having loop heights of 300 to 375 um survived more than 100,000 slow power cycles without a failure. In earlier

groups where the loop heights were only about 50 to 100 ym, failures began

to occur from a few hundred to a few thousand cycles. The bond separation in

all transistors was about 1.0 mm.

process dependent It is important to realize that the magnitude and effect of the flexingf3c tors induced stress is also affected by such factors as the extent of work hardening in the wire, deformation of the wire adjacent to the bond, and the existence of microcracks or other defects in the heel of the bond. That such

factors are important is indicated by the reports of Gaffney [68G1] where the

wire failure rate as a result of wire flexing resulting from low-repetitionrate pulsed-power operation was "severely" dependent on the bonding process

for aluminum wire bonds. Similarly, Villella and Nowakowski [70V1] , [71N3]

found the degree of deformation at the heel of the aluminum-wire bonds could

markedly affect the failures due to wire flexing, resulting from such on-off

power operation. Both Gaffney [69G1] and Villella and Nowakowski [70V1]

,

[71N3] noted that devices with gold-wire thermocompression ball or wedge

bonds did not exhibit this reliability problem, presumably because of the

greater ductility of gold. Phillips [71B5] has pointed out an additional

consideration when iron-nickel-cobalt headers are used. The differential

thermal coefficient of expansion of the header material and gold wire is

about half that for the case where aluminum wire is used. Thus, the thermally induced stresses in wire bonds made with gold wire would be significantly less than for wire bonds with aluminum wire, for such a header.

Recent work by Bevington et at. [70B1] with devices where the wire bonds are embedded in plastic serves as a good example to show that the test stress

depends greatly on the embedding materials used in the plastic devices.

Phenolic encapsulated devices were the most resistant to degradation due to thermal shock tests. Epoxy encapsulated devices were the least resistant

and silicone units were intermediate. The superior performance of the

phenolic encapsulated devices was attributed to the good match in the thermal

coefficient of expansion between the phenolic material and the lead frame

material.

There are two important questions about these test methods to which

is test n

degrading? there are no definite answers: (1) is the test degrading to the wire bonds

and (2) how many cycles constitutes a meaningful screening test? No general answers are available because they depend on the wire bonds and the associated device packages being tested, as indicated above, as well as on the

*F. Villella, Marshall Space Flight Center, Huntsville, Alabama 35812, private communication.

94
intended application of the device. With a better awareness of the factors

that can affect the magnitude of the stresses to wire bonds, appropriate

steps may be taken in the selection of constituent materials and in the

bonding processes and procedures to minimize such stresses. These steps

need to be taken at least in devices intended for use in such applications

as satellite transmissions and other telemetry services where the device

must sustain a great many on-off power cycles.

5.14. Electrical Tests

electrical tests

5.14.1. Introduction

A variety of electrical tests methods is in use but, aside from the

measurement of the bond interface resistance, little has been reported about

them in the literature. The use of an electrical test is attractive because

of the possibility of automating it, but no test has yet been found to be

sufficiently sensitive. Either the variable measured is insufficiently

sensitive to the kinds of defects which may be present, or the variable of

interest cannot be measured with sufficient sensitivity.

5.14.2. Bond Interface Resistance

The resistance of the bond interface is usually measured by a fourprobe method. In this method a small, known current is conducted through

the interface via one pair of contacts while the voltage across the interface is measured by a second pair of contacts.

The problem of insufficient sensitivity arises from the need to measure

a small contact resistance in series with the larger resistances of the wire

and metallization. Amlinger et at. [66A3] reported that the interface

resistance measured was primarily a function of wire size and material and was of the order of 1.0 mfi or less for gold or aluminum wire with diameter

from 0.7 to 2 mils. The resistance that is usually quoted includes contributions from the wire and film and is of the order of 100 mfi. An electrical

resistance of 100 mQ is equivalent to a length of about 1.5 mm of 1-mil

diameter gold or aluminum wire. The electrical resistivity of both materials

is about 3 x 10-6 fi-cm [70C1].

The contact resistance has not been found to be a satisfactory measure to judge the quality of the wire bond. Amlinger et at. [66A2] found that such

measurements did not give them insight into the mechanical strength of the

wire bond. In a continuation of the work, Riben et at. [67R1] could not

find any correlation between the contact resistance of the bond and the pull

strength of the wire bond and therefore concluded that the resistance could

not be used as a measure of mechanical integrity. However, because the pull

limitations

interface

resistance

sensitivity

mechanical

integrity

95
V-Icharacteristics

test is usually more a test of the wire rather than the bond adherence this

conclusion is expected. A more meaningful evaluation of the method would be

to determine the correlation which might exist between interface resistance

and shear strength.

degradation While Riben et al. [67R1] did not find the method capable of predicting

detection r t- & which wire bonds would fail as a result of process variations, they did believe it was useful in detecting the electrical degradation due to exposure

to elevated temperature of wire bonds with aluminum-gold interfaces.

clcctr i cs. X continuity 5.14.3. Electrical Continuity

The electrical continuity of the wire bond is usually tested by monitoring the voltage developed across the wire bond by a small test current.

A convenient method is to use a transistor curve tracer. Any irregularities

in the current-voltage curve or too great a voltage, as judged from previous

experience, are used as grounds for rejecting the wire bond. Usually these

criteria are symptoms of unsatisfactory electrical or physical continuity of

the bond. A number of laboratories use some variation of this empirical

test,

precautions Some precautions related to the use of the method on gold-aluminum

bonds have been suggested. Annular, micro-openings can form around gold ball

bonds to aluminum films as a result of the Kirkendall effect soon after

storage at as low a temperature as 125°C. To detect such defects may require

the applied voltage to be limited to millivolts to avoid the possibility of

arcing or other effects which would "heal" the micro-opening. Intermittent

open-circuits can occur after such a "healing" event [71L1]. Cunningham

suggested that the application of even a tenth of a volt could heal such

defects. The discharge of static charge has also been reported to heal microopenings [66B5], hence it is advisable to make these measurements in a static-charge-free environment.

To detect intermittent electrical discontinuities the test is made at temperatures sufficiently above and below room temperature so that thermal

temperature

effects expansion differences can perturb the wire bond. Browning et al. [66B5]

reported intermittent continuity problems as the measurement temperature was

changed. At one laboratory visited, the use of the test at an elevated

temperature (125°C) was considered important. This was due to experience

with bonds which passed the test at room temperature but exhibited intermittent electrical discontinuities at a higher temperature. At another

laboratory, the use of the test at below room temperature was recommended.

More recently, Brauer et al. [70B4] has recommended that in order to detect

*J. A. Cunningham, Texas Instruments, Stafford, Texas 77477, private communication.

96
intermittent bonds in plastic-encapsulated devices, electrical continuity

tests should be made at both the maximum and minimum rated temperature.

Haberer [71H2] has described this monitored thermal cycle test; its success

in detecting such defects has prompted the recommendation of this test for use in MIL-STD-883.

Mann [66M1] has described a circuit for automatically detecting those

wire bonds which exhibit too high a resistance. The circuit is designed

to register a defect when less than 7 yA is conducted for an applied voltage of 0.3 V. Other voltage-current settings may be used to improve the

detection of marginal bonds.

Another approach to the continuity test, suggested by Kessler [70B2],

is to use a high-current pulse of very short duration to detect poor bonds.

This approach is based on experience with high-current pulsing of laser

diodes. Weak bonds give an irregular current wave-shape pattern during

the pulse or separate at the wire-die interface.

monitored

thermal cycle

measurement

circuit

high-current

test

5.14.4. Noise Measurements

The generation of excess noise might be expected to result from the

presence of defects such as (1) a mechanical contact rather than a metallurgical bond between the wire and the bonding surface, (2) micro-cracks

at the heel of an ultrasonic bond, and (3) microvoids in the aluminum

metallization at the periphery of a gold wire ball bond. This would be

expected especially under a stress which would tend to move the wire bond

and thereby alter the current path. However, no reports have been found

which specifically deal with the use of noise measurements to detect weak

wire bonds. Maki et at. [65M1] have described noise measurements to detect defective electronic components, but no mention was made about

wire bonds.

5.15. Ultrasonic Bond Monitoring

no reports

US bond monitoring

5.15.1. Description

To identify weak ultrasonic bonds while they are being made, changes

in the coupling between the wire and metal film during the ultrasonic

bonding process are monitored. The monitoring has been attempted either

by detecting the ultrasonic energy transmitted through the bond and into

the substrate or by detecting the change in the loading of the ultrasonic

driving system.

Detecting the energy transmitted through the bond interface has been

somewhat less successful than detecting changes in loading. The earliest

attempt is one reported by Jones [62J1] who attached microphones to the

supporting base so that spot and seam bonding of thin aluminum sheets

two methods

detecting transmitted US energy

via microphones

97
via transducers

could be monitored. He reported that the tensile shear strength of the bonds

depended on the comparative ultrasonic energy transmitted through the bond.

Therefore, if the driving frequency drifted from its design or resonant

frequency a weaker bond would be made. The primary purpose of the monitor

was thus to assure that the driving frequency was at resonance. Later

Pruden and Schoenthaler [67P2] patented a method of determining the ultravia temperature- sonic energy dissipated in the substrate by detecting temperature-induced induced changes

changes in the electrical characteristics of the electrical component to which the wire bond is being made. In particular, they mentioned using the

voltage-current characteristics of an adjacent p-n junction. They maintained

that there exists a level of energy dissipation for which the bond quality

will be a maximum. Bond quality was judged in terms of "maximum strength

characteristics" and some minimum degree of bond deformation. It appears

that this approach may have some practical disadvantages because the optimum

level of energy dissipation measured, if it exists, would depend on such

things as the relative location and uniformity of response of the adjacent

temperature sensitive element monitored. In general it would not be possible

to have such an element near each bonding area on the die. Furthermore, such

a method would not be suitable for bonds made to the terminal.

More recently Bellin et at. [69B7] have explored the use of transducers

in the device holding fixture to measure the tangential forces during the

bonding cycle. The tangential force, f, is related to the shear modulus, M,

and the shear strain, S, of the aluminum wire, and the bonded area, A, by the

following expression

f = MAS or f = MAx/y,

where x is the lateral displacement of the top part of the wire with respect

to the bottom, and y is the thickness of the deformed wire. If the maximum

displacement of the wire is held constant by maintaining a constant driving

power, the monitoring transducer system will measure a quantity that is proportional to the ratio of the effective bond area to the bond thickness.

They proposed that this ratio is a function of bond adherence and could therefore be used as a measure to judge bond quality. Their work has not, however,

progressed to a point where they were able to substantiate this proposal nor determine at what point in time a good metallurgical bond is achieved. One

difficulty in interpreting the results of this technique may be related to recent observations by Leedy [70B3], She reported that bonding can occur at one or more sites within the area of the wire-metallization interface in an irregular and non-reproducible way from bond to bond (see figure 5). These

variations in the development and shape of bonded areas may affect the

tangential force measured and hence make interpretation difficult. Also,

98
ultrasonic stress waves can affect the physical properties of metals [66L3]

which implies that M will not be constant.

Several techniques have been explored for measuring the change in the

loading of the ultrasonic system during bonding. These methods have followed one of two general approaches: monitoring the displacement of the

transducer or bonding tool and monitoring the electrical parameters of

the ultrasonic driving circuit.

The first report found of monitoring displacement is by Worlton and

Walker [64W1] who described the use of a pickup cartridge in contact with

the tool to monitor the tool displacement during the bonding cycle. They

found that when the tool displacement amplitude decreased a good bond

would be made while if this decrease was not seen a weak bond would be

made. They also described a sensing circuit that stops the bonding cycle

after a predetermined decrease in tool displacement is detected.

A high-frequency microphone has been used to monitor the sound pressure from the tool and hence to measure qualitatively the tool displacement during bonding. The amplitude envelope of the sound pressure

decreases as a good bond is established. A detailed description of the

design and use of such a microphone system in addition to the use of a magnetic pickup is given by Harman and Kessler [71H1], Quantitative

measurements of the tool displacement have been made with laser inter

ferometric techniques [70W1], [71B1], [71B4].

Many commercial ultrasonic driving circuits incorporate feedback to minimize loading effects during loading. As a result, changes in the displacement amplitude of the transducer and tool are smaller. Therefore

the measurement of tool displacement for bond monitoring is made less

attractive because more sensitive detecting apparatus is required. An

alternative first suggested by Worlton and Walker [64W1] is to monitor

the electrical parameters of such driving circuits. Two commercial bonding machines incorporate such a bond monitoring approach.

Bonis* has described the use of small acceleration-detecting transducers attached to the transducer at various locations and to the top of

the bonding tool. He also suggested that, despite the experimental difficulties, measurements of the acoustic standing wave ratio in the transducer

could be used to monitor the impedance match at the bonding surface during

the bonding cycle. Martin^ has suggested that the various vibrational modes

of the tool, as might be detected by measuring the displacement at the top

of the tool with an attached accelerometer , might be relatable to the bondmaking process, but studies by Harman [70B2] of tool oscillations with a

detecting change

in loading

via pickup

cartridge

via microphone

via electronic

feedback

via acoustic SWR

*S. Bonis, Raytheon Company, Sudbury, Massachusetts 01776, private communication.

>C. T. Martin, ibid, private communication.

99
via harmonic

detection

discussion

high-resolution capacitor microphone indicate that the displacement at the

top of the tool does not necessarily relate to the displacement at the

bonding-end of the tool.

Martin* identified the third-harmonic component of the tool vibration

during bonding as a good indicator to use in bond monitoring. Later, in a computer simulation of the motion of a bonding tool, King [71K2] concluded

that monitoring the higher harmonic content of the tool acceleration would

be a more promising approach than monitoring either the acceleration magnitude (relatable to the tool displacement) or the acceleration phase. Harman

[71B4], [71B5], [72B2] has recently described a mixer system for monitoring

and studying harmonics of both oscillation amplitude and frequency during

bonding as detected with a high-frequency microphone. Significant increases

in the amplitude of the second and third harmonics during bonding were seen. These harmonics are best observed when the microphone is placed at a tool

node.

5.15.2. Discussion

needs development

sensitivity

The attraction for bond monitoring is the possibility of identifying

those isolated defective bonds that now go undetected and of detecting

process drift much sooner than is now practical. However, the methods for

monitoring the ultrasonic bonding process are still in an early stage of

development. It remains to be seen just how useful any of them will be.

Most of the methods explored so far depend on an interpretation of the shape

of the tool displacement envelope or some other measure of loading as a function of time. This is not desirable for production line testing because

of the additional time and judgment required for the interpretation.

Bonis"'" indicated that it was possible to identify strong bonds from the

change of loading during the machine bonding cycle. Furthermore, it was possible to identify weak bonds which a visual inspection would pass; in

these cases, the bond monitor indicated that either no coupling had occurred

or that decoupling had occurred during the bonding cycle. It appears, however, that bond monitoring may not be sufficiently sensitive to detect some

of the more subtle effects that may result in weak bonds. Under some conditions the monitor may indicate that a typical bond has been made, which a visual inspection will confirm, even though a weak bond has actually been

made. This has been observed when, for example, the cleanliness of the bonding surface or the adherence of the metal film to the substrate was

inadequate.

§

C. T. Martin, ibid, private communication.

+S. Bonis, ibid, private communication.

§H. K. Kessler and G. G. Harman, National Bureau of Standards, Washington, D. C. 20234,

unpublished results.

100
Bond monitoring, even if it lives up to all expectations, cannot be

considered as a complete screen for the wire bond. For example, such monitoring would not be expected to detect excessive weakening at the heel of

the first bond due to post-bond wire bending in the process of shifting to the second bonding site, nor would it necessarily detect excessive deformation or a crack at the heel of a bond.

Even if monitoring is not installed as a production line test, it can be useful in process control measurement and bonding process studies. Also,

it may be useful to apply the technique to other ultrasonic bonding schemes,

such as face-down bonding [71H1]

.

5.16. Correlation

limitations

process control

correlation

Few comments have been made regarding the correlation between different test methods for evaluating a given type of wire bond. Generally, no f ew reports

substantiating experimental results are offered in the literature and the

correlation or noncorr elation is dispensed with in a sentence or two. Those comments and results that have appeared in the literature are reviewed and evaluated below.

Generally, only a certain degree of correlation between pull and Pul1 and visual

inspection

visual inspection test can be expected. Rather than correlating with tests each other so that one test might be substituted for another, the two methods tend to supplement each other. A visual inspection test will re- ject wire bonds which a subsequent pull test might find acceptable.

Conversely, a visual inspection test will pass wire bonds which have

acceptable wire deformations but have essentially no adherence. Furthermore, the visual inspection test is concerned with eliminating potential

problems not directly related to pull strength, such as bond misplacement,

excessive loop height, etc. Recent work by Kashiwabara et at. [69K1] have shown some interesting bond deformation

and pull

correlations between the lateral bond deformation and the pull strength strength

of ultrasonic aluminum wire bonds. The pull strength was measured in a single-bond test where the direction of pull was 30 deg from the horizontal. They also found correlations between deformation and the ultrasonic bonding parameters of power, force, and time. All the bonds studied were made to aluminum films on an oxidized silicon substrate. They reported relative independence of the bond deformation on the aluminum film

thickness

.

To assist in establishing more meaningful deformation criteria, if

pull strength is to be the measure of quality, then more work of this kind

is needed. For example, similar information about the deformation of bonds

on the terminal and the effect of the condition of any metal film on the

101
pull, visual

Inspection, and electronic

anvil tests

pull and bond monitoring

tests

pull and centrifuge tests

terminal on the deformation would be extremely useful. It would also be

useful to determine what effect changes in the tensile strength of the wire

and the wire diameter would have on the correlation.

Bellin et al. [69B7] reported little correlation between the bond

deformation and pull strength of ultrasonic aluminum wire bonds. However,

the data shown to substantiate this claim, a plot of pull strength versus

bond deformation, is similar to that shown by Kashiwabara et al. [69K1]

except for more scatter in the points. It is not clear what causes this

increased scatter, but it may be due to the mixing of results of bonds made

with different bonding power and time settings.

Plough et al. [69P1] mentioned that the results of pull tests with the use of a visual inspection test could be used to predict the ability of

equivalent wire bonds to withstand the stresses imposed by an "electronic

anvil test" (referred to in section 5.12 as the short-duration-stress pulse

test)

.

Bellin et al. [69B7], in reporting some development work on a bond

monitoring technique, found little or no correlation between their pull

strength data and the outputs of their bond monitor. They questioned the use of pull strength data to judge completely the quality of the wire-tochip bond and suggested that tests such as those using ultrasonic energy

[69K4] be included in the evaluation of wire bonds. Their work only serves

to point out that, potentially, a bond monitor will tend to indicate the

quality of the bond itself while the pull test will provide an indication of

the strength of, primarily, the heel of the bond unless very weak bonds are

involved.

Good correlation between single-bond pull (at 30 deg from the horizontal)

and centrifuge (up to 60,000 g's) tests was claimed by Baker and Bryan [65B1] It is implied that this correlation was found for both aluminum and gold wire

bonds. Such correlation for aluminum wire would not be expected considering

the negligible stress that an acceleration of even up to 60,000 g's would

produce.

102
6. Summary summary

Two critical areas for reliability improvement of wire bonds are manufacturing process control and testing and evaluation methods. Despite the

variety of test methods used, no one or combination of methods has been

devised to cull out all wire bonds that will fail. In addition, the fabrication processes are either not sufficiently well understood or are not

sufficiently well controlled or both to be able to make the same wire bond

every time.

The essential features of the thermocompression and ultrasonic bonding

processes, the fabrication procedures, the bonding equipment, and the characteristics of the constituent materials of the wire bond pertinent to high

reliability are discussed in section 4. Also in section 4 is a review of the

degradation effects of the interaction of gold and aluminum in wire bonds

at elevated temperatures. It is emphasized that a large number of variables

must be considered and controlled when making wire bonds and that there is

relative ignorance of the quantitative (and in some cases qualitative) effects

that these variables have on the reliability of the wire bonds produced.

The methods used to test and evaluate wire bonds are discussed in section

5. In general there is little in the published literature that can be used

to evaluate the methods in terms of (1) the criteria to establish the reliability or quality of the wire bonds and (2) the effects of variations in

the test variables on the test results. An additional widespread problem

in the literature is the underspecif ication of both the methods used and the

wire bonds tested. Brief summaries of the comments made in section 5 about

the various methods are given in the following paragraphs.

The most widely used nondestructive (and nondegrading) test method for

wire bonds is the visual inspection test. It usually contains an extensive

list of reject criteria. The method can be used to detect relatively gross

defects but its success in detecting the more subtle defects is mixed.

The pull test is the most widely used destructive test. The method can be fast and easy to perform and it provides a number, the pull strength, for use as a quality measure. However, the temptation is to use only this number

without specifying other data needed to interpret the pull strength. Hence

pull-test data is often of limited value. A nondestructive form of the pull

test is also used. Skepticism that the method is actually nondegrading and

the need for more detailed reports of its effectiveness has hindered wide

use of the method.

The centrifuge test is also a widely used test. It is considered by

many to be useful for wire bonds with gold wire but not for those with

process control

test methods

fabr j cat ion variables

test criteria

and variables

visual

inspection

pull test

centrifuge test

103
air blast test

push test

mechanical shock

test

short-duration

stress pulse

test

variable fre- quency and vibration tests

ultrasonic

stress

temperature

cycling and thermal shock

electrical tests

aluminum wire because of the lower density of aluminum. Even for gold wire

it may be either inconvenient or damaging to other components of the device

if acceleration levels at which a significant stress can be imposed on the

wire bond are attained.

The air blast test is a relatively little used nondestructive test. An

air jet is directed at the wire bonds on the device to stress them mechanically. An often expressed reservation about the method is the difficulty

of quantifying the stress imposed on the wire bond. Consequently the test

is usually performed at a conservatively low air velocity to cull out only

the weakest wire bonds. The recently suggested push test, in which a probe

applies a deflecting force to the apex of the wire span, appears to offer,

in most cases, greater control over the applied stress but is slower because

each wire bond is tested individually. The push test method offers the

potential advantage over the nondestructive pull test in that the shape of

the wire span is not altered.

The mechanical shock test is generally regarded as being of marginal

value because of the small peak deceleration usually used in the test. One

limitation to increasing the stress is that the device package may be

damaged; providing the special jigging required to reduce package damage

is usually inconvenient.

The short-duration stress pulse test is in limited use. The tensile

stresses generated by the absorption of pulses of high-energy electrons can reach spall thresholds in the wire bond.

The variable frequency vibration and the vibration fatigue tests are of

marginal value, at best, because of the small stress imposed. The lowest

resonant frequency of most wire bonds is much higher than the maximum excitation frequency and the peak acceleration produces an essentially insignificant stress.

The ultrasonic stress method involving the use of focused ultrasonic

power to stress wire bonds has been explored empirically. Significant stress

levels can be reached with this method to cause failure in even strong wire

bonds. A better understanding of the nature of the stress imposed and how

its magnitude is affected by the test variables is needed before the method

can be evaluated.

Temperature cycling and thermal shock tests are widely used but there

has .been disagreement about the effectiveness of these methods. The stress

and hence the effectiveness of the tests is dependent on the differences in

the thermal coef f icients-of-expansion of the constituent parts of the device

and the shape of the wire span.

A variety of electrical test methods is in use to evaluate the electrical connection at and near the bond interface between the wire and the bonding surface. The use of an electrical test is attractive because of the

104
;

possibility of automating the test, but no test is sufficiently sensitive

to go much beyond detection of open circuits. Either the property measured

s insufficiently sensitive to the kinds of defects which may be present or the property cannot be measured with sufficient sensitivity.

The shear test is used to obtain a measure of the adhesion of gold shear test ball bonds. The method is not widely used and apparently suffers from

poor precision.

Bond monitoring methods are attractive for the case of ultrasonic ultrasonic bond monitoring bonding because they offer the possibility of checking each bond as it is

made to identify isolated defective bonds that otherwise could go undetected and of detecting unintentional variations in the fabrication process much sooner than is now practical. Various techniques for bond

monitoring are under consideration but all are in an early stage of

development and require much more work before adequate evaluations can be made.

105
7. APPENDIXES

7.1. Appendix A. Estimate of Lowest Resonant Frequency of Wire Bonds

To obtain an estimate for the lowest resonant frequency of a wire bond with a given

bond separation, assume that the wire loop describes a circular arc. That is, the wire

loop may approximate any circular arc from the extreme of a straight wire to that of a semicircular loop between the two bond sites. The three frequency modes to be considered are

extensional, non-extensional (node at mid-span), and lateral (perpendicular to the plane

of the wire loop). These modes are sketched in figure Al for the case where the wire describes a semicircular arc. Approximate values of the resonant frequencies of the lowest

extensional and non-extensional modes of vibration [28D1] and the lowest lateral mode of

vibration [34B1] have been calculated for clamped-end connections. The general expression

for the resonant frequency, f . , is given by,

(A.l) 01 1 ^ ,l ud

where E = modulus of elasticity (dynes/cm2 ), I = moment of inertia of area (cm1*),

p = mass per unit length of wire (g/cm)

,

d « separation of bonds (cm) , and

C. = frequency coefficient of the mode.

i The frequency coefficients for the three modes, C. , C_, and C_, are all dependent on

the subtending angle, 0. In addition, C. also depends on the ratio of the bending to torsional stiffness, k, and C_ also depends on the value of R = (d/r) [ (d/4h) + (h/d) ]

, where

r is the wire radius. The variations of C. for each of the three modes as a function of the

i subtending angle and the ratio of bond separation to loop height, d/h, are shown in figure

Al. The values of C.. are for the case where the ratio of the bending to torsional stiffness,

k, is equal to two.i" The dependence of C_ is given for two values of R. As may be seen

from figure Al and equation (A.l), the lateral mode of vibration will have a significantly

lower resonant frequency than either of the other two modes, except for small subtending

angles where the resonant frequencies of the lateral and extensional modes approach each

other.

*The product of the subtending angle and the radius of curvature of the wire arc is equal

to the wire length.

fk » EI/RJ for a circular cross-section, where E and R are, respectively, the modulus of

elasticity and of rigidity, and I and J are, respectively, the moment and the polar moment

of inertia of area. The value of k = 2 was taken as a convenient upper-bound. This was done with the recognition that Ci is not a sensitive function of k within the range of

values reported for aluminum and gold [68S3]. The use of a value of k = 1.25, which would

be a lower bound for the values of k reported, would increase the value of C^ by less than

3 percent.

106
VIBRATION MODES

)[- R=lOO #£V. C3-NONj£fhJr? EXTENS.ONAL

79 -rr//^////

JL

R\ C2" EXTENSIONAL

Ci- LATERAL

TV

= 180°

40 80° 120°

e

160°

J L_J ' 40 20 I 10 8 7 6 S 4

15 d/h

Figure Al. Variation of the frequency coefficients, C^, with subtending angle, 0, and ratio

of bond separation to maximum wire span height, d/h, for each of three vibration modes,

identified by the sketched modes for = 180 deg. The coefficient C2 depends on the value

of R = (d/r) [(d/4h) + (h/d)], where r is the wire radius. The dependence of C2 on is

given for two values of R.

107
The dependence on bond separation of the resonant frequency for the lateral vibration

mode is shown in figure A2 for 1-mil diameter gold and aluminum wire bonds with a circular

wire loop for three different values for d/h: 2, 5, and infinity. For d/h = 2, the wire

loop is a semicircle, a highly unlikely configuration. It is expected that most wire loops

will lie between the arc where d/h = 5 and a straight wire (d/h infinite) . These calculations are based on the assumption that the wire ends are clamped to the bond sites. In real

wire bonds the condition of the stiffness of the wire connection to the bonding site should

be between that for a clamped and for a hinged, free-roll connection. Hence for more con- servative lower bounds for the resonant frequency of wire bonds, the frequency coefficient,

C , and consequently all curves in figure Al, may be lowered by about 25 percent to allow

for a less stiff connection.

'f . was calculated using the values for E of 7.0 x 10 11 dynes/cm2 and 7.8 x 10 11 dynes/cm2 for aluminum and gold, respectively.

108
00

60

40

20h

N

* 10-

4

2

CURVES

No. I-AI (d/h)=®

2-AI (d/h)=5

3-Au (d/h)=co

4-Au (d/h)*5;AI (d/h)«2

5-Au (d/h)= 2

.02 .06 JO .14

d (cm)

.18 .22 .26

Figure A2. Dependence of the resonant frequency, fQ ^, on bond separation, d, for the lateral

mode of vibration for 1-mil diameter gold and aluminum wire bonds with circular shaped wire

spans having the ratios of d/h equal to 2 (semicircular), 5, and infinity (straight wire).

The height of the span at its apex is h.

109
7.2. Appendix B. Expression Relating Four Measures for Pull Rate

Four measures of pull rate in the pull test are mentioned in section 5.3.2: rate of

straining, R ; rate of stressing, R ; elapsed time for completing part or all of the test,

t; and speed of the moving head during the test, v. These measures are related below for

the simple case of the wire bond pulled at mid-span and in a direction perpendicular to the

bonding surface as shown in figure Bl.

e zlz eE £F

T % — % % — £ (B.l)

R vh2 R 2AR h n s s

where R = rate of straining, per unit gauge length (s-1

),

R = rate of stressing, per unit area (N cm

-2

s-1

),

t elapsed time in pull test (s)

,

v = speed of pull, per unit gauge length (s_1

),

e = wire elongation,

% « wire distance from bond to mid-span,

h wire loop height,

E = Young's modulus of elasticity (N cm

-2

), and

F = pull strength of wire bond (N)

.

The above relations may be obtained as follows: First, R = (1/SL) (dJL/dt) % (l/£)

(A£/At) » e/x, where t is time; therefore t £ e/R . Second, v = (1/h) (dh/dt) £ (1/h)

(Ah/At); but dh - (£/h)d£ and Ah £ (£/h)A£, so v £ U/h2

) or (Afc/At) % vh2 /£. As already

seen, x = e£/(AS./At) ; therefore x £ e£ 2 /vh2

. Third, R = R /E. Hence, from the first relation, x ^ eE/R . Fourth, R - (1/A) (dF /dt) , where F is the tensile force in the wire and

' s » s w w A is the cross-sectional area. But, F = F£/2h. Therefore R % (£/2Ah) (AF/At) ^ (£/2Ah) w s (F/x), or x % «.F/2AhR . s

*For aluminum and gold, E is, respectively, about 7.0 x 10 6 N/cm2 and 7.8 * 106 N/cm2

.

110
Figure Bl. Illustration of double-bond pull test.

Ill
8. REFERENCES

8.1. Introduction

identification

code

ordering

availability

addresses

guidance

location in text

abbreviations

Each citation has been given an identification code which consists of

a sequence of two digits, a letter, and another digit. The first two digits

indicate the year of publication and the letter is the initial of the first

author's surname. The last digit is used to distinguish those papers which

would otherwise have the same code. No rule was used in the assignment of

the last digit.

The papers in the bibliography are arranged according to their codes.

The codes are grouped first by year, then in alphabetical order by letter,

and then in numerical order by the last digit.

Some citations are followed by notes which refer to additional information intended to assist in obtaining the referenced work. Reports available

from the National Technical Information Service (NTIS) , Sills Building,

5285 Port Royal Road, Springfield, Virginia 22151, are followed by a number

preceded by the letters AD or PB, or the letter N. This is the NTIS Accession number which should be used when ordering.

A number of other citations, generally to conference papers, are followed by a number either in brackets or parentheses. The number refers

to one of the addresses listed in section 8.3. If the number is in brackets

the address listed is one to which an order may be placed for the paper or the conference proceedings. If the number is in parentheses the address is

that of the first author's place of work at the time the paper was published.

Some citations are also followed by guidance in brackets. For example,

reference may be made to the pages in the paper that are relevant to the

subject.

All citations are followed by the page number (s) in the text where the

paper is cited. These page numbers are in braces.

The journal or conference abbreviations generally follow those of the

Chemical Abstracts. In order to minimize any possible confusion, those

journals abbreviated are listed in alphabetical order by their abbreviations

in section 8.4. Additional abbreviations are included which are used in

citations to some conference meetings.

The identification codes used are taken from a comprehensive bibliography on wire bonds

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67A1 Arleth, J. M. , and R. D. Demenus

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6801 Ogilvie, R. E., S. H. Moll, and M. A. Schippert

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{pp. 12, 35}

6852 Sullivan, P. 69B2

PAUL SULLIVAN SPEAKS OUT ON FLIP

CHIP VS. BEAM LEAD, EEE, vol. 16,

pp. 60-63, Dec. 1968. {p. 2}

69A1

6 9A3

69B1

Samsonov, G. V., Ed.

HANDBOOK OF THE PHYSICOCHEMICAL

PROPERTIES OF THE ELEMENTS, IFI/

Plenum, New York, 1968. [translated

from Russian] {p. 106}

Tanaka, S., and K. Chiba

SEMICONDUCTOR DEVICE UTILIZING AN

AUAL2 LAYER AS A DIFFUSION BARRIER

THAT PREVENTS 'PURPLE PLAGUE', U. S. Patent 3,401,316; Sept. 10, 1968.

{p. 22}

Uthe, P. M. THE SOLID STATE WELD, Uthe Technology,

Inc. Technical Newsletter, vol. 1,

May 1968. [summarized in 69U1] (5)

{p. 32}

Uthe, P. M. THE FRICTION OF NON-LUBRICATED METALS,

Uthe Technology, Inc. Technical Newsletter, vol. 1, June 1968. [summarized in 69U1] (5) {p. 32}

Uthe, P. M. WELDING, Uthe Technology, Inc. Technical Newsletter, vol. 1, July

1968. [summarized in 69U1] (5)

{pp. 32, 37}

Uthe, P. M. A SIMPLE WIRE BONDER, Uthe Technology,

Inc. Technical Newsletter, vol. 1,

Aug. 1968. [summarized in 69U1] (5)

{p. 18}

1969

Ang, C. Y., P. H. Eisenberg, and

H. C. Mattraw

PHYSICS OF CONTROL OF ELECTRONIC

DEVICES, Proc. 1969 Annual Symp. on Reliability, Chicago, 111., pp. 73-

85, Jan. 1969. [see pp. 76, 82]

(13) {p. 72}

Antler, M. WHAT DO GOLD PLATING SPECS REALLY

MEAN?, Products Finishing, vol. 34, pp. 56-66, Oct. 1969. {p. 16}

Bell, J. L. UPGRADING OF MICROELECTRONIC TEST

PROCEDURES FOR MILITARY HI-REL

ACHIEVEMENT, Trans. 23rd Annual Technical Conf., pp. 767-770, Los Angeles,

Calif., sponsored by American Society

for Quality Control, Ann Arbor, Mich.,

May 1969. {p. 93}

Binelli, W. D., and R. H. Soltau

DEVELOPMENT OF QUALIFICATION TEST

PROGRAM FOR MICROELECTRONIC DEVICES,

Final Rpt. (Nov. 1, 1968 to July 3,

1969), Contract No. NAS 1-8714,

Sept. 1969. N70-11544 {pp. 77, 82}

117
69B3 Budd, J. B. DIE AND WIRE BONDING CAPABILITIES OF

REPRESENTATIVE THICK-FILM CONDUCTORS,

Solid-State Technol., vol. 12, pp. 59-63, June 1969.. {pp. 16, 28, 32}

69B5 Bullis, W. M. , Ed.

METHODS OF MEASUREMENT FOR SEMICONDUCTOR MATERIALS, PROCESS CONTROL,

AND DEVICES, NBS. Technical Note 488,

Quarterly Rpt. (Jan. 1 to March 31,

1969), July 1969. [see pp. 21-25]

[30] (29) {pp. 15, 58, 59}

69B6 Bullis, W. M. , Ed.

METHOD OF MEASUREMENT FOR SEMICONDUCTOR MATERIALS, PROCESS CONTROL,

AND DEVICES, NBS Technical Note 495,

Quarterly Rpt. (April 1 to June 30,

1969), Sept. 1969. [see pp. 24-31]

[30] (29) {pp. 41, 42}

69B7 Bellin, J. L. S., A. E. Brown, A. S. Hamamoto, and G. C. Knollman

PIEZOELECTRIC MONITOR OF MICROELECTRONIC WIRE BONDS, Lockheed Rpt.

LMSC B-62-69-9, June 1969. (21)

{pp. 98, 102}

69B8 Brinton, J. MIL STD 883 — A REAL TEST CASE,

Electronics, vol. 42, pp. 131-136,

Aug. 18, 1969. {p. 50}

69C2 Cline, J. E. , J. M. Morris, and

S. Schwartz

SCANNING ELECTRON MIRROR MICROSCOPY

AND SCANNING ELECTRON MICROSCOPY OF

INTEGRATED CIRCUITS, IEEE Trans.

Electron Devices, vol. ED-16, pp. 371-375, April 1969. {p. 56}

69D1 Dudderar, T. D. THE EFFECT OF GRIP STRESSES ON THE OCCURRENCE OF FAILURE IN TENSION

TESTS OF WIRE, Mater. Res. Std.,

vol. 9, pp. 26-30, Oct. 1969.

{p. 10}

69D3 MILITARY SPECIFICATION MICROCIRCUITS

GENERAL SPECIFICATION FOR —, MilM-38510, Nov. 20, 1969. [1] {p. 50}

69G1 Gurland, J. MICROSTRUCTURAL ASPECTS OF THE

STRENGTH AND HARDNESS OF CEMENTED

TUNGSTEN CARBIDE, Contract No. SD-86,

Dec. 1969. AD 699187 {p. 94}

69G2 Grable, R. C, and H. E. Patzer

WIRE BONDING APPARATUS FOR MICROELECTRONIC COMPONENTS, U. S. Patent

3,430,835; March 4, 1969. {pp. 11,

25}

69H1 Heinen, K. G., and G. B. Larrabee

THE DETERMINATION OF RESIDUAL PHOTORESIST ON SILICON USING RADIOTRACER

IODINE-131, Solid-State Technol.,

vol. 12, pp. 44-47, April 1969.

{p. 19}

69H2 Hamiter, L. C, Jr. HOW RELIABLE ARE MOS IC'S? AS GOOD

AS BIPOLARS, SAYS NASA, Electronics,

vol. 42, pp. 106-110, June 23, 1969.

{pp. 2, 19}

69K1 Kashiwabara, M., S. Nakayama, and

M. Suzuki

SETTING AND EVALUATION OF ULTRASONIC

BONDING FOR AL WIRE, Rev. Elec.

Commun. Lab., vol. 17, pp. 1014-1021,

Sept. 1969. {pp. 38, 57, 58, 59, 101,

102}

69K2 Krieg, R. D., and W. B. Murfin

STRUCTURAL CONSIDERATIONS IN ELECTRONIC MICROCIRCUIT LEAD WIRES, March 1969. PB 183544 {pp. 84, 89}

69K3 Kashiwabara, M. , and S. Hattori

FORMATION OF AL-AU INTERMETALLIC COMPOUNDS AND RESISTANCE INCREASE FOR

ULTRASONIC AL WIRE BONDING, Rev. Elec.

Commun. Lab., vol. 17, pp. 1001-1013,

Sept. 1969. {pp. 17, 21, 22}

69K4 Knollman, G. C, A. S. Hamamoto, and

J. L. S. Bellin

REPORTS ON ULTRASONIC SCREENING OF

TRANSISTORS AND INTEGRATED CIRCUITS,

Lockheed Rpt. LMSC B-62-69-8, June

1969. (21) {pp. 76, 102}

69L1 Lauf fenburger , H. A., and T. R. Myers

SUMMARY AND INTERPRETATION OF RELIABILITY DATA ON VARIOUS MICROCIRCUIT

BONDING TECHNIQUES, Proc. Holm Seminar

on Electrical Contact Phenomena, pp. 61-68, Nov. 1969. [6] {pp. 2, 73}

69L2 Laub, J. L., and M. N. Mansour

WIRE CLAMP, U. S. Patent 3,430,834;

March 4, 1969. {p. 49}

6901 0'Connell, E. P. AN INTRODUCTION TO MIL-STD-883 TEST

METHODS AND PROCEDURES FOR MICROELECTRONICS, Proc. 8th Reliability

and Maintainability Conf., Denver,

Colorado, pp. 530-542, July 1969.

[10] (27) {pp. 2, 20, 50, 52, 56,

77, 78, 80, 88}

6902 Ono, K. , M. Nishihata, and S. Kobayashi

FINE ALUMINUM TRANSISTOR LEAD WIRES,

Rev. Elec. Commun. Lab., vol. 17, pp. 974-988, Sept. 1969. {p. 10}

69P1 Plough, C, D. Davis,' and H. Lawler

HIGH RELIABILITY ALUMINUM WIRE BONDING, Proc. Electronic Components

Conf., Washington, D. C, pp. 157-

165, April-May 1969. [2] (28)

118
66P1 (continued)

{pp. 10, 13, 14, 17, 19, 31, 38, 47,

48, 49, 56, 74, 87, 93, 102}

6951 Slemmons, J. W. THE MICROWORLD OF JOINING TECHNOLOGY,

American Welding Society 50th Annual

Meeting and Welding Exposition;

Philadelphia, Pa., April-May 1969.

[8] (13) {pp. 22, 23, 27, 35, 56,

72}

6952 Schnable, G. L., and R. S. Keen

ALUMINUM METALLIZATION ADVANTAGES

AND LIMITATIONS FOR INTEGRATED CIRCUIT APPLICATIONS, Proc. IEEE, vol.

57, pp. 1570-1580, Sept. 1969.

{p. 15}

6953 Selikson, B. VOID FORMATION FAILURE MECHANISMS IN INTEGRATED CIRCUITS, Proc. IEEE, vol.

57, pp. 1594-1598, Sept. 1969.

{pp. 15, 19, 20}

6954 Shurtleff, W. O. RELIABILITY HANDBOOK FOR SILICON

MONOLITHIC MICROCIRCUITS VOLUME 2 —

FAILURE MECHANISMS OF MONOLITHIC

MICROCIRCUITS, Contract No. NAS 8-

20639, April 1969. N69-23226 [see

pp. 2-IV-2 to 2-IV-6] {pp. 56, 77,

81, 82, 85, 86}

69T2 Tarowsky, N. HOW TO ASSEMBLE HYBRID MICROWAVE IC'S,

Microwaves, vol. 8, pp. 52-62, Aug.

1969. {p. 22}

69U1 Uthe, P. M. VARIABLES AFFECTING WELD QUALITY IN ULTRASONIC ALUMINUM WIRE BONDING,

Solid-State Technol., vol. 12, pp. 72-76, Aug. 1969. {pp. 12, 14, 31,

34, 37, 40, 47}

69U2 Uthe, P. M., Jr., L. G. Wright, and

R. E. Greenan

ULTRASONIC FREQUENCY POWER SUPPLY,

U. S. Patent 3,445,750; May 20, 1969.

{p. 42}

69W1 Watkins, L. S. INSPECTION OF INTEGRATED CIRCUIT

PHOTOMASKS WITH INTENSITY SPATIAL

FILTERS, Proc. IEEE, vol. 57, pp. 1634-1639, Sept. 1969. {p. 53}

69W2 Watkins, L. S. INTEGRATED CIRCUIT PHOTOMASKS INSPECTION BY SPATIAL FILTERING, SolidState Technol., vol. 12, pp. 29-32,

Sept. 1969. {p. 53}

1970

70A1 Adams, M. A. AN INVESTIGATION OF THE STRENGTH OF

ALUMINUM WIRE USED IN INTEGRATED

CIRCUITS, NASA Tech. Brief 70-10275,

Aug. 1970. [9] {p. 58}

70A3 STANDARD SPECIFICATION FOR GOLD WIRE

FOR SEMICONDUCTOR LEAD-BONDING, (ASTM

DESIGNATION: F72-69) , 1970 Annual

Book of ASTM Standards, part 8, 1970.

[11] {pp. 11, 12}

70A4 STANDARD METHODS OF TESTING FINE ROUND

AND FLAT WIRE FOR ELECTRON DEVICE AND

LAMPS, (ASTM DESIGNATION: F219-67)

,

1970 Annual Book of ASTM Standards,

part 8, 1970. [11] {p. 10}

70A5 STANDARD METHOD FOR MEASURING

DIAMETER OF FINE WIRE BY WEIGHING,

(ASTM DESIGNATION: F205-63) , 1970

Annual Book of ASTM Standards, part

8, 1970. [11] {p. 10}

70A6 STANDARD METHODS OF TENSION TESTING

OF METALLIC MATERIALS, (ASTM

DESIGNATION: E8-69) , 1970 Annual

Book of ASTM Standards, part 31,

1970. [11] {pp. 10, 71}

70A7 STANDARD METHOD OF TEST FOR RESISTIVITY OF ELECTRICAL CONDUCTOR

MATERIALS, (ASTM DESIGNATION: B193-

65), 1970 Annual Book of ASTM

Standards, part 8, 1970. [11]

{p. 10}

70A8 STANDARD SPECIFICATION FOR IRONNICKEL-COBALT SEALING ALLOY, (ASTM

DESIGNATION: F15-69), 1970 Annual

Book of ASTM Standards, part 8, 1970.

[11] { P . 16}

70B1 Bevington, J. R. , J. P. Cook, and

D. R. Little

PLASTIC IC RELIABILITY EVALUATION AND

ANALYSIS, 8th Annual Proc. Reliability Physics, IEEE Catalog No. 70C59-PHY, pp. 73-80, 1970. [2]

{p. 94}

70B2 Bullis, W. M. , Ed.

METHODS OF MEASUREMENT FOR SEMICONDUCTOR MATERIALS, PROCESS CONTROL,

AND DEVICES, NBS Technical Note 520,

Quarterly Rpt. (July 1 to Sept. 30,

1969), March 1970. [see pp. 32-43]

[30] (29) {pp. 23, 41, 43, 44, 45,

46, 97, 99}

70B3 Bullis, W. M. , Ed.

METHODS OF MEASUREMENT FOR SEMICONDUCTOR MATERIALS, PROCESS CONTROL,

AND DEVICES, NBS Technical Note 527,

Quarterly Rpt. (Oct. 1 to Dec. 31,

119
70B3

70B4

70B5

70B6

70B7

70B8

70C1

70D1

70D2

(continued)

1969), May 1970. [see pp. 31-47]

[30] (29) {pp. 17, 33, 41, 42, 44,

49, 98}

Brauer, J. B., V. C. Kapfer, and

A. L. Tamburrino

CAN PLASTIC ENCAPSULATED MICROCIRCUITS PROVIDE REALIABILITY WITH

ECONOMY?, 8th Annual Proc. Reliability Physics, IEEE Catalog No. 70C59-PHY, pp. 61-72, 1970. [2]

(27) {p. 96}

Bradfield, G. ULTRASONIC TRANSDUCERS — 1. INTRODUCTION TO ULTRASONIC TRANSDUCERS,

PART A, Ultrasonics, vol. 8, pp. 112-

123, April 1970. {p. 43}

Bradfield, G. ULTRASONIC TRANSDUCERS — 1. INTRODUCTION TO ULTRASONIC TRANSDUCERS,

PART B, Ultrasonics, vol. 8, pp. 177-

189, July 1970. {p. 43}

Bullis, W. M. , and A. J. Baroody, Jr.,

Eds.

METHODS OF MEASUREMENT FOR SEMICONDUCTOR MATERIALS, PROCESS CONTROL,

AND DEVICES, NBS Technical Note 555,

Quarterly Rpt. (Jan. 1 to March 31,

1970), Sept. 1970. [see pp. 27-36]

[30] (29) {pp. 49, 60}

Bullis, W. M. and A. J. Baroody, Jr.,

Eds.

METHODS OF MEASUREMENT FOR SEMICONDUCTOR MATERIALS, PROCESS CONTROL,

AND DEVICES, NBS Technical Note 560,

Quarterly Rpt. (April 1 to June 30,

1970), Nov. 1970. [see pp. 29-38]

[30] (29) {pp. 32, 34, 38, 59}

Cox, W. P., E. E. Anderson, and

J. H. Anderson, Jr. ULTRASONIC ALUMINUM WIRE BONDING FOR

MICROELECTRONIC APPLICATIONS, Proc.

1970 Annual Symp. on Reliability,

Los Angeles, Calif., vol. 3, pp. 228-

236, Feb. 1970. (31) {pp. 11, 13,

14, 70, 95}

Devaney, J. R. APPLICATION OF SCANNING ELECTRON

MICROSCOPY TO INTEGRATED CIRCUIT

FAILURE, Solid-State Technol., vol.

13, pp. 73-77, March 1970. {p. 12}

Davis, D. FACTORS IN HIGH RELIABILITY WIRE

BONDING, 8th Annual Proc. Reliability Physics, IEEE Catalog No. 70C59-PHY, pp. 170-176, 1970.

[similar to 69P1] [2] (28) {pp.

14, 17, 18, 38, 40, 45, 49}

70D3 MILITARY STANDARD — TEST METHODS

FOR SEMICONDUCTOR DEVICES, Military

Standard 750B, Feb. 27, 1970. [1]

{pp. 50, 77, 81, 82, 85, 86, 88}

70E1 MINICOMPONENTS - BUILDING BLOCKS FOR

HYBRIDS, EDN, vol. 15, pp. 33-44,

Sept. 1, 1970. [see p. 43] {p. 35}

70M3 McCormick, J. E. IC PACKAGING: THE VARIATIONS,

Electronic Packaging and Production,

vol. 10, pp. 234-238, Jan. 1970.

{p. 2}

70P1 Pankratz, J. M. , and D. R. Collins

A COMPARISON OF 1% MG-AL AND 1% SI-AL

WIRE INTERCONNECTS, 8th Annual Proc.

Reliability Physics, IEEE Catalog No. 70C59-PHY, pp. 163-169, 1970. [also

published in IEEE Trans. Reliability,

vol. R-19, pp. 89-94, Aug. 1970] [2]

(32) {pp. 13, 14, 38, 69}

70P2 Philofsky, E. INTERMETALLIC FORMATION IN GOLDALUMINUM SYSTEMS, Solid-State

Electron., vol. 13, pp. 1391-1399,

Oct. 1970. [also 8th Annual Proc. Reliability Physics Symp., pp. 177-

185, 1970] {pp. 19, 21}

70S1 HIGH-PRESSURE PROCESS MAKES WIRE BY

SQUEEZING, Spectrum, vol. 7, pp. 21,

Aug. 1970. {p. 10}

70V1 Villella, F., and M. F. Nowakowski

INVESTIGATION OF FATIGUE PROBLEM IN 1-MIL DIAMETER THERMOCOMPRESSION AND

ULTRASONIC BONDING OF ALUMINUM WIRE,

NASA Technical Memorandum, NASA TM

X-64566, pp. 1-45, Nov. 30, 1970.

N71-16494 {p 94}

70W1 Wilson, A. D., B. D. Martin, and

D. H. Strope

HOLOGRAPHIC INTERFEROMETRY APPLIED

TO MOTION STUDIES OF ULTRASONIC

BONDERS, IEEE Ultrasonics Symp.,

San Francisco, Calif., Oct. 21-23,

1970. [2] (34) {p. 99}

1971

71B1 Bullis, W. M., Ed.

METHODS OF MEASUREMENT FOR SEMICONDUCTOR MATERIALS, PROCESS CONTROL,

AND DEVICES, NBS Technical Note 571,

Quarterly Rpt. (July 1 to Sept. 30,

1970), April 1971. [see pp. 23-32]

[30] (29) {pp. 43, 63, 99}

71B2 Boylan, J. R. THERMOCOMPRESSION BONDING, IEEE

Intern. Conv. Digest, New York,

Session 5B, pp. 598-599, March 1971.

[2] {pp. 2, 26, 35}

120
71B4 Bullis, W. M., Ed.

METHODS OF MEASUREMENT FOR SEMICONDUCTOR MATERIALS, PROCESS CONTROL,

AND DEVICES, NBS Technical Note 592,

Quarterly Rpt. (Oct. 1 to Dec. 31,

1970), Aug. 1971. [see pp. 34-45]

[30] (29) {pp. 11, 58, 63, 71, 99,

100}

71B5 Bullis, W. M., Ed.

METHODS OF MEASUREMENT FOR SEMICONDUCTOR MATERIALS, PROCESS CONTROL,

AND DEVICES, NBS Technical Note 598,

Quarterly Rpt. (Jan. 1 to March 31,

1971), Oct. 1971. [see pp. 20-28]

[30] (29) {pp. 38, 41, 93, 94, 100}

71G1 Glass, R. A., T. G. Maple, and R. D. Wales

INTERCONNTECTION PROBLEM AREAS IN MICROCIRCUITS, IEEE Intern. Conv.

Digest, New York, Session 5B, pp. 248-249, March 1971. [2] {pp. 35,

48}

71G2 Goldfarb, S. WIRE BONDS ON THICK FILM CONDUCTORS,

Proc. Electronic Components Conf.,

Washington, D. C, pp. 295-302, May

1971. [2] (36) {p. 16}

71H1 Harman, G. G., and H. K. Kessler

APPLICATION OF CAPACITOR MICROPHONES

AND MAGNETIC PICKUPS TO TUNING AND

TROUBLE SHOOTING OF MICROELECTRONIC

ULTRASONIC BONDING EQUIPMENT, NBS

Technical Note 573, pp. 1-22, May

1971. {pp. 40, 41, 49, 99, 101}

71H2 Haberer, J. R. TECHNIQUES FOR DETECTING STRESS

INDUCED INTERMITTENT FAILURES IN ENCAPSULATED DEVICES, IEEE Intern.

Conv. Digest, New York, Session 7CJ,

pp. 612-613, March 1971. {pp. 88,

97}

71H3 Hart, R. R. A WIRE EXTENSOMETER FOR DETERMINING

THE MECHANICAL PROPERTIES OF FINE

WIRES, Mater. Res. Std., vol. 11,

pp. 26-28, April 1971. {p. 10}

71J1 Johannes en, F. ULTRASONIC ALUMINUM WIRE BONDING,

IEEE Intern. Conv. Digest, New York,

Session 7CI, pp. 600-601, March 1971.

{pp. 18, 40, 43, 48, 49}

71J2 Joshi, K. C. THE FORMATION OF ULTRASONIC BONDS

BETWEEN METALS, Welding J., vol. 50, pp. 840-848, Dec. 1971. {pp. 2, 31,

32, 35, 43, 70}

71K1 Kalvelage, B. F. A PNEUMATIC SHOCK TESTER FOR ELECTRON DEVICES, Solid-State Technol.,

vol. 14, pp. {p. 84}

57-59, March 1971.

71K2 King, CM.

DYNAMIC SIMULATION OF AN ULTRASONIC

WIRE BONDING TOOL, Raytheon Report

ER71-4135, Contracts N0003070-C0055

and N0003071-C0061, Jan. 25, 1971.

(35) {p. 100}

71L1 Leyshon, W. E., and R. E. Warr

AN OVERVIEW OF HYBRID INTEGRATED

CIRCUIT RELIABILITY PROBLEMS AND SOLUTIONS, IEEE Intern. Conv. Digest,

New York, Session 7CJ, pp. 606-607,

March 1971. {pp. 12, 96}

71M1 Matcovich, T. J. INTERCONNECTIONS IN HYBRID CIRCUITS,

IEEE Intern. Conv. Digest, New York,

Session 5B, pp. 240-241, March 1971.

{p. 2}

71N1 LINE CERTIFICATION REQUIREMENTS FOR

MICROCIRCUITS, NHB 5300. 4(3C), May

1971. [30] {pp. 50, 52, 56}

71N2 TEST STANDARDS FOR MICROCIRCUITS,

NHB 5300. 4(3D), May 1971. [30]

{pp. 58, 59, 70}

71N3 Nowakowski, M. F., and F. Villella

THERMAL EXCURSION CAN CAUSE BOND

PROBLEMS, 9th Annual Proc. Reliability Physics Symp. IEEE Catalog

No. 71-C-9-PHY, pp. 172-177, 1971.

[2] (39) {p. 94}

71P1 Philofsky, E. DESIGN LIMITS WHEN USING GOLDALUMINUM BONDS, 9th Annual Proc. Reliability Physics Symp., Las Vegas,

IEEE Catalog No. 71-C-9-PHY, 1971.

[2] (37) {pp. 17, 18, 22}

71P3 Polcari, S. M. , and J. J. Bowe

EVALUATION OF NONDESTRUCTIVE TENSILE

TESTING, Report No. DOT-TSC-NASA-71-

10, pp. 1-46, June 1971. N71-37516

{p. 72}

71R1 Ravi, K. V., and E. Philofsky

THE STRUCTURE AND MECHANICAL PROPERTIES OF FINE DIAMETER ALUMINUM —

1 PCT SI WIRE, Metallurgical Transactions, vol. 2, pp. 711-717, March

1971. {pp. 11, 13}

71U1 Uthe, P. M., Ed.

Uthe Technology International Technical Newsletter, vol. 2, No. 5, Dec. 1971. (5) {p. 2}

1972

72B1 Bullis, W. M., Ed.

METHODS OF MEASUREMENT FOR SEMICONDUCTOR MATERIALS, PROCESS CONTROL,

121
72B1 (continued)

AND DEVICES, NBS Technical Note 717,

Quarterly Rpt. (July 1 to Sept. 30,

1971), April 1972. [see pp. 22-30]

[30] (29) {pp. 15, 63, 89, 94}

72B2 Bullis, W. M. , Ed.

METHODS OF MEASUREMENTS FOR SEMICONDUCTOR MATERIALS, PROCESS CONTROL,

AND DEVICES, NBS Technical Note 727,

Quarterly Rpt. (Oct. 1 to Dec. 31,

1971), June 1972. [see pp. 29-45]

[30] (29) {pp. 15, 34, 37, 63, 94,

100}

72B3 Bullis, W. M. , Ed.

METHODS OF MEASUREMENT FOR SEMICONDUCTOR MATERIALS, PROCESS CONTROL,

AND DEVICES, NBS Technical Note 733,

Quarterly Rpt. (Jan. 1 to March 31,

1972), August 1972. [see pp. 27-33]

[30] (29) {pp. 2, 89}

72H1 Harman, G. G. , and K. 0. Leedy

AN EXPERIMENTAL MODEL OF THE MICR0_

ELECTRONIC ULTRASONIC WIRE BONDING

MECHANISM, Presentation Abstracts,

Reliability Physics Symposium,

IEEE Catalog No. 72-CH0692-4-PHY,

April 1972. [2] (29) {pp. 19, 32,

35}

72H2 Horsting, C. W. PURPLE PLAGUE AND GOLD PURITY,

Presentation Abstracts, Reliability

Physics Symposium, IEEE Catalog No. 72-CH0692-4-PHY, April 1972. [2]

(36) {p. 19}

72R1 Ravi, K. V., and E. M. Philofsky

RELIABILITY IMPROVEMENT OF WIRE BONDS

SUBJECTED TO FATIGUE STRESSES,

Presentation Abstracts, Reliability

Physics Symposium, IEEE Catalog No. 72-CH0692-4-PHY, April 1972. [2]

(37) {pp. 14, 89}

72S1 Schafft, H. A. WIRE-BOND ELECTRICAL CONNECTIONS:

TESTING, FABRICATION AND DEGRADATION — A BIBLIOGRAPHY 1957-1971,

NBS Technical Note 593, Jan. 1972.

[30] (29) {pp. 5, 112}

122
8.3. Addresses

[1] U. S. Naval Publications and Forms Ctr., 5801 Tabor Avenue, Philadelphia, Pennsylvania

19120.

[2] Publications Sales Dept., The IEEE, 345 East 47th Street, New York, New York 10017.

[3] American National Standards Institute, Inc., 1430 Broadway, New York, New York 10018.

[4] Industrial and Scientific Conference Management, Inc., 222 West Adams Street, Chicago,

Illinois 60606.

Uthe Technology International, 320 Soquel Way, Sunnyvale, California 94086.

Illinois Institute of Technology, Chicago, Illinois 60606.

ASME Order Dept., 345 47th Street, New York, New York 10017.

American Welding Society, 345 East 47th Street, New York, New York 10017.

Technology Utilization Division, NASA, Code UT, Washington, D. C. 20546.

Gordon and Breach, Science Publishers, Inc., New York, New York 10001.

ASTM, 1916 Race Street, Philadelphia, Pennsylvania 19103.

MIT Instrumentation Laboratories, Cambridge, Massachusetts 02142.

North American Aviation/Autonetics, Anaheim, California 92803.

Norden Division of United Aircraft Corp., Norwalk, Connecticut.

Philco-Ford Corp., Microelectronics Div., Blue Bell, Pennsylvania 19422.

Westinghouse Electric Corp., Baltimore, Maryland.

Weltek Div., Wells Electronics Inc., South Bend, Indiana.

Aeroprojects Inc., West Chester, Pennsylvania.

Weldmatic Div., Unitek Corp., Monrovia, California.

Martin Marietta Corp., Quality Engineering Dept., Orlando, Florida.

Lockheed Missiles and Space Co., Lockheed Palo Alto Res. Labs., Palo Alto, California

94304.

Sandia Labs., Albuquerque, New Mexico 87115.

Hughes Aircraft Co., Welder Dept., Oceanside, California 92054.

United Aircraft Corp., Electronic Components Div., Trevose, Pennsylvania.

Westinghouse Defense and Space, Mfg., Research and Development, Baltimore, Maryland.

McDonnell Douglas Astronautics Co., Western Div., Santa Monica, California.

RADC, Griff iss Air Force Base, Rome, New York 13440.

Fairchild Semiconductor, Research and Development Labs., Palo Alto, California 94304.

National Bureau of Standards, Washington, D. C. 20234.

Superintendent of Documents, U. S. Government Printing Office, Washington, D. C. 20402,

Lockheed Missiles and Space Co., Sunnyvale, California.

Texas Instruments, Inc., Dallas, Texas 75222.

AC Electronics Div., General Motors Corp., Milwaukee, Wisconsin.

IBM Corp., Endicott, New York 13760.

Raytheon Co., Sudbury, Massachusetts 01776.

RCA, Somerville, New Jersey 08876.

Motorola Inc., Phoenix, Arizona 85008.

Delco Electronics Division, General Motors Corp., Milwaukee, Wisconsin.

NASA Marshall Flight Center, Huntsville, Alabama 35812.

123
8.4. Abbreviations

AIME - American Institute of Mining, Metallurgical and Petroleum Engineers

ASME - American Society of Mechanical Engineers

ASTM - American Society for Testing and Materials

ASTME - American Society of Tool and Manufacturing Engineers

Appl. Phys. Lett. - Applied Physics Letters

Bell Lab. Rec. - Bell Laboratories Record

Brit. Commun. Electron. - British Communication and Electronics

Brit. J. Appl. Phys. - British Journal of Applied Physics

Circuits Mfg. - Circuits Manufacturing

Conf. - Conference

EDN - (formerly Electrical Design News)

EE - (now, The Electronic Engineer; formerly, Electronic Industries)

Electron. Packag. Prod. - Electronic Packaging and Production

Electron. Prod. - Electronic Products Magazine

Electronic Engineer - The Electronic Engineer (formerly EE, formerly Electronic Industries)

Electron. Lett. - Electronics Letters

Electro-technol. - Electro-technology (New York)

Eval. Eng. - Evaluation Engineering

IEEE - Institute of Electrical and Electronics Engineers (formerly IRE)

IEEE Intern. Conv. Record - IEEE International Convention Record (formerly IRE ...) IEEE Trans. Electron Devices - IEEE Transactions on Electron Devices (formerly IRE ...) IEEE Trans. Nucl. Sci. - IEEE Transactions on Nuclear Science

IEEE Trans. Pts. Materials Packaging - IEEE Transactions on Parts, Materials, and Packaging

IEEE Trans. Sonics Ultrason. - IEEE Transactions on Sonics and Ultrasonics

Int. - International

IEC - International Electrotechnical Commission

IRE - Institute of Radio Engineers

IRE Intern. Conv. Record - IRE International Convention Record

IRE Trans. Electron Devices - IRE Transactions on Electron Devices

J. Appl. Phys. - Journal of Applied Physics

J. Electrochem. Soc. - Journal of the Electrochemical Society

J. Sci. Instrum. - Journal of Scientific Instruments

Mater. Eval. - Materials Evaluation

Mater. Res. Std. - Materials Research and Standards

Metals Eng. Quart. - Metals Engineering Quarterly

Mfg. - Manufacturing

Nat. Electron. Conf. - National Electronics Conference

NEPCON - National Electronic Packaging and Production Conference

Philips Tech. Rev. - Philips Technical Review

Proc. IEEE - Proceedings of the Institute of Electrical and Electronics Engineers

Prod. Eng. - Product Engineering

124
Rev. Sci. Instr. - The Review of Scientific Instruments

Rev. Elec. Commun. Lab. - Review of the Electrical Communications Laboratory. Tokyo.

(Denki Tsushin Kenkyujo)

SAE - Society of Automotive Engineers

Semicond. Prod. - Semiconductor Products

Semicond. Prod. Solid State Technol. - Semiconductor Products and Solid State Technology

Soc. - Society

Solid State Electron. - Solid State Electronics

Solid State Technol. - Solid State Technology

Symp. - Symposium

Technol. - Technology

Trans. Met. Soc. AIME - Transactions of the Metallurgical Society of the AIME

Welding J. - Welding Journal

WESCON - Western Electric Show and Convention

125
SUBJECT INDEX

Air blast test, 74, 75

Adherence

of bonds (tests for), 58, 59, 73-76, 86,

87, 97 of gold films, 15, 73, 74

Degradation and failure modes

of device, 12

of wire, 12-14

of wire bond, 3, 17, 18, 20-22, 27, 48,

49, 74, 89, 90, 96

B Ball bonds, 6, 23, 24, 28, 29 Beam lead bonding, 2 Bond adherence (see Adherence)

Bond deformation, 6, 10, 11, 15, 17-19, 24,

28, 48, 54, 57, 101, 102

Bond efficiency, 70

Bond interface resistance test, 95, 96

Bond schedule (see also Thermocompression

bonding and Ultrasonic bonding)

effect of wire, 14, 15, 28

effect of surface contamination, 19

Bonding surface, 15-19

aluminum, 15-17

sintering temperature, 17 coefficient of friction, 28

factors affecting bonding, 16-19

factors affecting later degradation, 17,

18

gold, 15

plating specifications, 16

gold/molybdenum, 15, 74 hardness (aluminum) , 17 intermetallic (AuAl£) , 22

multilayer films, 15

roughness, 18, 19

thermocompression bonding, 16

thickness (aluminum, gold), 17, 18, 21

thick film, 16, 17, 28

firing temperature, 16

Electrical continuity test, 96, 97

monitored thermal cycle test, 97

high-current stress test, 97

Electrical resistance of wire bond, 95

Electrical tests, 95-97

Electronic anvil test, 87

Eyelet bond, 26, 27, 59

Face-down bonding, 2, 101

Flyer-plate test, 87

Gold aluminum interactions, 19-22

intermetallic compounds, 16, 19-22

Kirkendall voids, 16, 17, 20, 21, 96

H High-current stress test, 97

I Intermetallic compounds (see also Gold

aluminum interactions), 16, 19-22

K Kirkendall voids (see also Gold aluminum

interactions), 16, 17, 20, 21, 96

Centrifuge test, 4, 5, 76-82, 102

test condition, 76, 77 analysis, 78-80

evaluation, 80

Correlation, 95, 101, 102

pull strength and bond deformation, 101,

102

pull test and ultrasonic bond monitoring,

102

pull and centrifuge tests, 102

pull and contact resistance, 95

pull and visual inspection, 101

pull, electronic anvil, and visual

inspection, 102

Lift-off patterns (of ultrasonic bonds), 32-

34

M

Mechanical shock test, 81-84

test conditions, 81, 82

evaluation, 81

stress (of the test), 83, 84

Metallization (see Bonding surface)

MIL-STD-750B, 50

MIL-STD-883, 50

126
Monitored thermal cycle test, 88, 96, 97

N Noise measurements, 97

of wire bonds, 83, 84, 106-108

Ribbon wire, 15, 34, 37 Rigidity (see Ultrasonic bonding and Ultrasonic bonding equipment)

Order of testing, 77, 80,

Plague (see Gold aluminum interactions)

Plate-slap test, 87

Power cycling test, 18, 88, 89, 94, 95

Pull strength (see Pull test)

Pull test, 57-73, 101, 102

analysis (resolution of forces) , 60-68

bond efficiency, 70

double-bond pull, 57 evaluation of test, 69

evaluation of wire bond, 38, 39

failure modes, 58, 59 gauge (gram-force), 70, 71

nondestructive test, 57, 59, 72, 73

pull strength, 57 correlation with bond deformation,

101, 102

degradation (thermally induced) , 21 dependence on bonding schedule, 38,

39

dependence on pull rate, 71

dependence on shape of wire loop, 63,

66, 69, 70

pulling probe, 59, 71

location on wire loop, 72

rate of pull, 59, 71 110

single-bond pull, 57, 59, 60

specification needs, 69-72 use in

evaluating metallizations for thermocompression bonding, 16

measuring degradation effects, 17, 21

measuring effects of fabricating processes and procedures, 17

developing bonding schedules, 38

Push test, 75, 76

Rate of pull (see Pull test)

Reliability of wire bonds, 2, 4, 5 Resistance (of wire bond) , 95

increase after high temperature exposure, 17, 21

Resonant frequency

of terminal posts, 18

related to ultrasonic bonding machines,

40-42

Shear test, 73, 74 Short-duration stress pulse test, 86, 87,

102

Spider bonding, 2 Stress to wire bond

deflection, 83, 89

differential thermal expansion, 89-93

flexing, 89-93

mechanical shock, 83

tensile, 89

tensile stress wave, 86

vibration, 83

wire fatigue, 74

Temperature cycling test, 87-95

analysis (of wire flexing) , 89-93

degradation effects, 18, 94, 95

evaluation, 93

monitored, 88

stress of test, 89, 93

Temperature effects on ultrasonic bonding, 41, 42

wire, 10, 12-14

wire bond, 17, 21, 22, 96, 97 Terminal (see also Bonding surface)

base material, 16

gold plating

finish, 18, 19

specifications, 16

metal plating, 16, 18

reduction in height, 18

Test methods (general), 50, 51

Testing (order of), 77, 80, 88

Thermal shock test (see Temperature cycling

test)

Thermocompression bonding, 22-31

ball bond, 23-25

ball size, 28-30

bonding schedule, 27, 28

coefficient of friction, 28 eyelet bond, 26, 27 flame-off procedure, 23, 26, 29

force (bonding) , 28 hot gas technique, 23, 26, 29 intimate contact, 22

mechanisms , 22

procedure for making wire bond, 23, 25

stitch bond, 23-26

127
Thermocompression bonding (continued)

temperature (bonding)

methods for attaining, 29 used, 22, 28, 29

thick film (bonding to), 16, 17, 28

time (bonding), 28, 29 versus ultrasonic bonding, 27, 35

wire alignment, 24, 26, 30

wire removal techniques, 23, 25

yield point, 23

Thick films, 16, 17, 28

Thin films (see Bonding surface)

Tool (for thermocompression bonding)

design, 26, 30

finish, 30

material, 30, 45 wear, 29, 30

Tool (for ultrasonic bonding)

anchoring, 45

bonding schedule (effect on) , 45

design, 45, 47, 48

drop velocity, 48, 49

extension, 45, 46

finish, 47

mechanical characteristics (importance

of), 45

tipping, 48

vibration

amplitude, 41, 43

modes, 41-44

Tuning (see Ultrasonic bonding equipment)

U Ultrasonic bond monitoring, 97-101, 102

methods, 97-100

evaluation, 100, 101

Ultrasonic bonding (see also Ultrasonic

bonding equipment and Tool (for ultrasonic bonding) ) , 31-49

alignment, 35, 37

bonding schedule, 34, 35, 37, 38, 45

coupling of ultrasonic energy, 35

force (bonding) , 37

heights of bonding surfaces, 18, 35

intimate contact, 31, 34, 35

limitations, 35, 37

machine requirements, 39, 40

mechanisms, 31, 32

melt hypothesis, 31, 32, 34, 35

metallographic examinations, 32

lift-off patterns, 32-34

power (bonding) , 37 procedure for making wire bonds, 35-37

rigidity (need for) , 35

silicon precipitates, 32, 34

temperature measurements, 32

time (bonding) , 37

unwanted movement, vibration, or energy

absorption, 39, 40, 43, 45, 49

versus thermocompression bonding, 27, 35

weld formation, 32-34

with aluminum wire, 34, 35

with gold wire, 35

Ultrasonic bonding equipment {see also Tool

(for ultrasonic bonding)), 39-43, 48, 49 adjustment, 40

components (principal) , 40

device chuck and stage, 49

diagnosing problems, 40, 43

movement (unwanted), 39, 40, 43, 49

operation speed, 48, 49

Q of electro-mechanical system, 40, 41, 45

transducer and coupler assembly, 43

mechanical stage, 48

tuning, 40-42

ultrasonic energy absorption (unwanted)

,

43

ultrasonic power supply, 40-42

tuning, 41, 42

effect of temperature, 41

effect of power supply impedance, 41

frequency sweep modes, 42

feedback stabilization, 42

wire clamp, 49

wire flexing, 48

Ultrasonic cleaning bath, 76

Ultrasonic energy (effect on wire plasticity)

,

31

Ultrasonic stress test, 76

Units, 9

Variable frequency vibration test, 84, 85

Vibration fatigue test, 85, 86

Visual inspection, 52-57, 101, 102

limitations, 56-57

reject criteria, 52-56

SEM examination, 54

test philosophies, 52

x-ray examination, 54

W Wire, 10-15

alignment, 15, 24, 26, 30

aluminum

1% magnesium 10, 14, 34

degradation, 14

1% silicon, 10, 12-14

degradation, 12-14

silicon precipitates, 12-14', 32, 34

annealing, 12-14

bond schedule (affected by), 14, 15

characteristics (measurement of), 10, 11

contamination, 12, 14

dimensions, 10, 11, 28

drawing, 10

128
Wire (continued)

ductility, 13

elongation, 11

fatigue, 14

gold, 10, 35

handling care, 12, 15

hardening, 10

hardness, 10, 28

lubricants, 12, 14

plasticity, 31

resistivity, 10

ribbon, 15, 34, 37

specifications, 11, 13

spool for wire, 12, 30

stabilized (thermodynamically) , 14

surface, 11

temperature effects, 12-14

tensile strength 10-14

yield point, 23

Wire bond

components, 2, 3 failure points, 3, 27, 28, 38

importance of, 2 procedure for fabricating by

a thermo compress ion process, 23, 25 an ultrasonic process, 35-37

reliability, 2, 4 resonant frequencies, 106-108

129
FORM NBS-114A (1-71) 2. Gov't Accession No.

U.S. DEPT. OF COMM. BIBLIOGRAPHIC DATA

SHEET

1. PUBLICATION OR REPORT NO. NBS TN-7263. Recipient's Accession No.

4. TITLE AND SUBTITLE Testing and Fabrication of Wire-Bond Electrical Connections

A Comprehensive Survey

5. Publication Date September 1972

6. Performing Organization Code

7. AUTHOR(S)

Harry A, Schafft

8. Performing Organization

10. Project/Task/Work Unit No. 4251522 & 4251533

9. PERFORMING ORGANIZATION NAME AND ADDRESS

NATIONAL BUREAU OF STANDARDS

DEPARTMENT OF COMMERCE

WASHINGTON, D.C. 20234

11. Contract/Grant No.

E07 2-813 & P0-2-0054

12. Sponsoring Organization Name and Address Dept. of Navy

Strategic Systems Project Office

Washington, D. C. 20360

(continued below)

13. Type of Report & Period Covered Final

14. Sponsoring Agency Code

15. SUPPLEMENTARY NOTES

also supported by:

Defense Nuclear Agency

Wpgh-mgl-rm, T) . C. . 20305

16. ABSTRACT (A 200-word or less factual summary of most significant information. If document includes a significant bibliography or literature survey, mention it here.) The fabrication and testing of wire-bond electrical connections used in integrated circuits, hybrid circuits, and low-power discrete semiconductor devices are surveyed comprehensively. The survey is generally restricted to wire-bond electrical connections where the wire diameter is less than 2 mils and where the wire

is bonded either by thermocompressive or ultrasonic means. Under the general

heading of fabrication, the essential features of the thermocompression and ultrasonic bonding processes, the fabrication procedures, and the characteristics of

the constituent materials of the wire bond pertinent to high reliability are surveyed. Also included is a review of the interaction of gold and aluminum as one

of the primary failure mechanisms in wire bonds. Both new and old test methods

are surveyed with emphasis on their capabilities and limitations. In particular,

the following test methods are discussed: visual inspection; pull, shear, air

blast, push, ultrasonic stress, centrifuge, mechanical shock, variable frequency

vibration, vibration fatigue, short-duration stress pulse, temperature cycling,

thermal shock, bond interface resistance, and electrical continuity tests; noise

measurement; and ultrasonic bond monitoring. Analyses of some of the methods

with regard to the stress that the test imposes on the wire bond have been made

and the results are used in discussing the relevant methods.

Key words (continued): microelectronics; reliability; survey (wire bond); semiconductor

devices; testing (wire bond); thermocompression bonding; ultrasonic bonding; wire bond.

17. KEY WORDS (Alphabetical order, separated by semicolons) Bonding; degradation (wire bond); discrete devices; electrical interconnection; fabrication (wire bond); failure (wire bond); hybrid circuits; integrated circuits; (cont.)

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