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NBS SPECIAL PUBLICATION 400-70

U.S. DEPARTMENT OF COMMERCE / National Bureau of Standards
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NATIONAL aVBlAV

OV tTAKDARDV

LUmAlIT Semiconductor Measurement Technology:

The Use of Acoustic Emission to Determine

ttie Integrity of Large Kovar Glass-Sealed

Microelectronic Packages

George G. Harman

Semiconductor Devices and Circuits Division

National Bureau of Standards

Washington, DC 20234

Sponsored by Naval Avionics Center

Indianapolis, IN 46218

U.S. DEPARTMENT OF COMMERCE, Malcolm Baldrige, Secretary

NATIONAL BUREAU OF STANDARDS, Ernest Ambler, Director

Issued May 1982

JUL 1 9 1982
Library of Congress Catalog Card Number: 82-600528

National Bureau of Standards Special Publication 400-70

Nat. Bur. Stand. (U.S.), Spec. Publ. 400-70, 80 pages (May 1982) CODEN: XNBSAV

U.S. GOVERNMENT PRINTING OFFICE

WASHINGTON: 1982
Table of Contents

Page

Abstract , 1

Introduction 1

Part A Vibration Experiments 4

A-1 The Effects of Avionics Environmental Vibration on the Seals of

PC Board-Mounted Hybrid Packages 4

A-1.1 Introduction 4

A-1. 2 Theoretical Considerations , 5 A-1 .3 PC Board Layout Considerations 8

A-1. 4 Experimental Measurements 9

A-1. 4.1 The Equipment Setup and Preliminary Vibration Tests of PC

Boards 9

A-1. 5 Preliminary Experiments on Vibration of PC Board-Mounted

Hybrid Packages 13

A-1. 6 The Vibration Characteristics of Metallization-Patterned PC

Boards with Soldered-In Hybrid Packages 14

A-1. 7 Susceptibility of Plug-in-Type Packages to Fatigue Damage at

High-g Levels 17

A-1. 8 Long-Term Low-g Nonresonant Vibration Effects 21

A-1. 9 Effect of Very High-g Forces on Flatpacks 21 A-2 Acoustic-Emission Equipment and Techniques for Detecting Glass

Seal Cracking During Vibration 23

A-2.1 Development of an Acoustic-Emission Detector for Use in

Vibration Experiments 23

A-2. 2 Determining AE Waveform Signatures in Order to Separate

Glass Cracking From Metal Fatigue 25

A-3 Acoustic-Emission-Monitored Vibration Experiment 30

A-3.1 Equipment Setup 30

A-3. 2 Conclusions of the Vibration Experiments 36

Part B Nondestructive Stress Tests for Hybrid Package Integrity ... 38

B-1 Leak Test Measurements and Procedures 38

B-1.1 Development of a High-Temperature Open-Package Leak Test

Procedure 38

B-1. 2 Preconditioning Procedure for Hybrid Packages Before Hermetic

Testing 38

B-2 Development of an AE-Monitored Thermal Shock Test for Open

Hybrid Packages 41 B-2.1 Introduction 41

B-2. 2 Characterization of the Hot Stage Shock Test 42

B-2. 3 Acoustic-Emission-Monitored Shock Test 44

B-2. 4 The Possible Use of a Thermoelectric Temperature Cycler for

Stressing Hybrid Packages in Conjunction with AcousticEmission Monitoring 51

B-3 Determination of the Susceptibility of Kovar Glass Seals to

Damage During Wire Bonding 52

iii
Page

B-r3.1 Mechanical Characteristics of Kovar Leads in Hybrid Packages . 52

B-3.2 Characteristics of Kovar Leads under Simulated Wire Bonding

Stresses . , 55

B-3.3 Acoustic-Emission-Monitored Tests to Determine Glass-Seal

Damage During Bonding 58

B-3.3.1 Room-Temperature-Bonding Experiments 58

B-3.4 AE-Monitored Package Lead Bending under Thermosonic Bonding

Conditions 62

B-3.5 Metallurgical Properties of Hybrid Package Leads 66

B-4 Conclusions of Part B 67

Acknowledgments 70

References 71;

List of Figures

1 . Sketch of "fixed" boundary conditions 5

2. Sketch of "supported" boundary conditions ... 5

3. Sketch identifying dimensions of a vibrating plate 6

4. Pictorial examples of the first several vibration modes of rec- tangular boards 7

5. An overview of the vibration equipment 10

6. The sinusoidal vibration requirements for equipment designed for

installation in jet airplanes 11

7. The resonant vibration peaks on a 20-min frequency sweep of an unloaded PC board 12

8. Photomicrographs of a 1-in. square hybrid package after vibration

testing 15

9. A closeup. of the vibration mount and the printed circuit board

with three flatpack hybrids soldered on it 16

10. Acceleration versus frequency for an etched PC board similar to

that in figure 9, but containing no hybrids 18

11. Acceleration versus frequency for a PC board similar to that in

figure 9 except that it contains one center-mounted flatpack

hybrid 19

12. Acceleration versus frequency for a PC board similar to that

shown in figure 12 except that three plug-in-type hybrids were mounted on the board 20

iv
Page

13. SEM photomicrographs of glass-metal seals in hybrid packages

that have been vibrated 22

14. Steps in the assembly of the TO-100 acoustic-emission detectorpreamplifier 24

15. The frequency spectral response of a TO-100 acoustic-emission

detector-preamplifier (

1 -MHz peak response) 26

16. The frequency spectral response of a TO-100 acoustic-emission

detector-preamplifier (5-MHz peak response) 27

17. The frequency spectral response of a high-quality commercial

AE detector 28

18. Fast Fourier transform of an acoustic-emission waveform obtained

by intentionally bending a lead on a hybrid until the glass seal

cracked 29

19. Block diagram of equipment setup for AE monitoring and control

of the PC-board hybrid seal experiment 31

20. A printed circuit board with three plug-in-type hybrids soldered

to it 3 3

21. An enlargement of the leftmost hybrid in figure 20 showing more details of the TO-100 AE detector 34

22. Composite of the acceleration experienced by a flatpack hybrid

and the acoustic-emission burst emitted at the peak acceleration

of the solid curve 35

23. The high-temperature open package leak test apparatus 39

24. Temperature-rise curves for a hybrid package on the hot stage of

figure 25 43

25. Thermal shock hot stage with a 1- by 1-in. hybrid package and a water-cooled AE detector on top 46

26. Block diagram of general circuitry used for acoustic-emission

tests and data analysis 47

27. Cumulative acoustic-emission counts versus time for the AEmonitored hot stage shock test (in units of 100) 48

28. Cumulative acoustic-emission counts versus time for the AEmonitored hot stage shock test (in units of 1000) 49

29. Acoustic-emission bursts resulting from stressing Bi-Te boules . 53

V
Page

30. Sketch identifying beam-bending variables 54

31 . A force apparatus used to simulate bonding forces on hybrid leads 56

32. An expansion of the probe and sample portion of figure 31 ... . 56

33. Acoustic-emission burst from glass cracking due to a force of

45 gf applied to the internal lead 59

34. A large acoustic-emission burst resulting from a small glass

sliver cracking off from the bottom side of the lead and becoming

a loose particle 59

35. An AE burst at 25 gf applied to the lead from a seal that was damaged on the outside with no damage observed on the inner side

of the seal 61

36. Apparatus to apply thermosonic-type forces and monitor any seal

cracking in hybrid packages 63

37. The first 10 percent of a 50-ms ultrasonic bonding pulse as detected by the AE detector in the equipment setup of figure 36 . 64

38. A 5X expansion of the waveform in figure 37 centered on the position of the arrow 64

39. The initial 8 percent of a second 60-kHz bonding pulse applied to

the same bond that was completely formed by the pulse shown in

figure 37 . 65

40. Photomicrographs of kovar leads from six different hybrid

packages 68

List of Tables

1 . Calculated Modal Frequencies for Two PC Boards 8

2. Comparison Between Calculated and Measured Modal Frequencies for an Unloaded Board that was Copper Clad on One Side 13

3. Measured Thicknesses of Hybrid Package Leads 55

4. Comparison Between Measured and Calculated Bending of Kovar Leads

for Leads of 0.01 x 0.014 x 0.10 in 57

5. Kovar Lead Bending Corrected for Glass Elasticity 58

vi
Semiconductor Measurement Technology

The Use of Acoustic Emission to Determine the Integrity of Large

Kovar Glass-Sealed Microelectronic Packages

by

George G. Harman

Semiconductor Devices and Circuits Division

National Bureau of Standards

Washington, DC 20234

Abstract

The general objective of this research was to develop tests to determine the integrity of large hybrid packages under various thermal and mechanical stresses that may be encountered during assembly, during installation in systems, or in operation. Several mea- surement techniques were investigated, but emphasis was placed on acoustic-emission test procedures. The accomplishments were: (1)

The effects of avionics environmental vibration on the seals of

hybrid packages mounted on printed-circui t (PC) boards were determined. A major conclusion of this section was that lead fatigue

failure occurs before seal damage on packages from high quality

lots. (2) A small acoustic-emission detector was developed that is

sensitive to surface waves, but relatively insensitive to vibrationinduced cable noise. (3) A high-temperature (125°C) open-package

helium leak test method was successfully developed to observe marginal seal damage. (4) An acoustic-emission test for inspection of

hybrid packages during high-temperature thermal shock was developed. (5) A study of possible damage to seals during thermocompression and thermosonic bonding, during lead forming, and during

other assembly operations was carried out. A general conclusion of

this study is that the glass-metal seals in packages from known

high quality lots are very reliable even when subjected to high

stresses. However, the seals from packages "screened as good" from

reject or poor quality lots are subject to hermetic failure under

moderate stresses. There is little correlation between visual

inspection failures of glass seals and their hermeticity.

Key words: acoustic emission; hermeticity; hybrid microelectronics; hybrid packages; microelectronic packaging; thermal shock;

vibration.

INTRODUCTION

The general objective of this study was to develop tests to determine the

integrity of large hybrid packages subjected to various thermal and mechanical stresses that they may encounter during assembly or later during operation. Several measurement techniques were investigated, but emphasis was placed on acoustic-emission test procedures. A preliminary report on the

work was published in June 1980 [1]. Parts of that report included here

serve as background or necessary introduction to the final acoustic-emission

evaluations

.

1
Acoustic emission (AE) is generally defined as being a transient elastic wave

or stress wave generated by the rapid release of energy within a material

when that material undergoes fracture or deformation. The classic example of

this has been known for years as "tin cry" where merely bending a piece of

tin will result in an audible sound. In 1950, Kaiser made the first comprehensive study of the phenomena [2]. His name is associated with the generally irreversible nature of AE, in which little or no acoustic emission occurs

until previously applied stress levels are exceeded. The emitted stress

waves may have frequencies ranging from the audible into the megahertz region, but the maximum energy is usually concentrated in the mechanical resonance modes of the test specimen. Detection of these waves usually takes

place with ceramic piezoelectric transducers that are acoustically coupled to

the specimen; however, wide band optical and capacitive detection methods

have also been used. Publications are available that give theory, equipment,

and applications of AE to a variety of nonelectronic applications [3,4].

The various sources of acoustic emission that have been observed include:

crack nucleation and propagation; twinning; grain boundary sliding; multiple

dislocation slip; creation of multiple dislocations; solid-solid, solidliquid, and liquid-solid phase transformations; and the Barkhausen effect

(realignment of magnetic domains). Most microelectronic uses of AE are con- cerned with solid-liquid-solid phase transformations (as in welding or soldering) or with crack propagation (as when a weld or a glass seal breaks).

However, recently there have been AE studies of dislocation propagation in

gallium arsenide LEDs and in silicon wafers.

The wavelength of AE in the usual detection frequency range is typically on the order of a centimeter. In the case of nuclear pressure vessels and other

large structures, this wavelength offers sufficient resolution for triangulation to determine the source location of the AE. However, for electronicsized samples multiple internal reflections inevitably mask the position of

the AE source, and one must infer this from the position of the applied

stress that caused the emission. Investigators are usually constrained to

attach a transducer, which may be larger than the specimen, in whatever manner is possible (such as using tapered acoustic waveguides), and work with

whatever signal is received. Waveform signatures, of frequency and amplitude, are recorded and empirically correlated with appropriate mechanical

stress te.sts (e.g., destructive bond pull tests) for interpretation. The

interpretation of AE signals from typical electronics applications may never

be amenable to clear mathematical solution.

Recently, acoustic emission has been used to evaluate electronic materials

and assembly processes. The largest effort in this area took place at the

Western Electric Engineering Research Center, Princeton, New Jersey [5]. The

first published use of AE in electronics was by Vahaviolos [6] . He used AE

to reveal substrate cracking during the thermocompression bonding of beam

lead semiconductor devices to gold metallization on ceramic substrates.

Saifi and Vahaviolos [7] have reported the use of AE for real-time nondestructive evaluation of laser spot welding of small polymer insulated wires

to electronic terminal posts. Carlos and Jon [8] used AE to detect cracks

in the ceramic casings of capacitors that resulted from the thermal shock of

soldering. Cracks so generated could not be observed during visual inspection because the crack-susceptible area was covered with solder. Jon e-t at.

2
[9,10] have described the use of AE for the nondestructive evaluation of the

quality of several types of complex-metallurgy resistance welds such as those

used in welding leads to tantalum capacitors. Another evaluation of resistance welding quality by AE monitoring of small diameter nickel wire welding

in electronic assemblies was carried out by Knollman and Weaver [11]. Ikoma

et aZ. [12] have used acoustic-emission measurements to investigate the con- ditions that produce crystallographic damage in GaAs LEDs during thermocompression wire bonding. Harman [13] used AE for nondestructive evaluation of

beam lead, flip-chip, TAB, and hybrid chip capacitor bonds.

3
PART A VIBRATION EXPERIMENTS*

A-1 The Effects of Avionics Environmental Vibration on the Seals of PC

Board-Mounted Hybrid Packages

A-1 .Introduction

There have been a number of studies in the past to determine the effects of

vibration on printed circuit (PC) board-mounted components. These have dealt

primarily with such electronic components as resistors, capacitors, and transistors. In all of these studies, metal fatigue of the leads resulting from

printed circuit board vibration was assessed. No studies have been carried

out to determine more subtle damage such as vibration-induced cracks that may

be produced and propagated in the glass hermetic seals of hybrids or other

susceptible components mounted on the circuit board. If the hermetic seals

were damaged, eventually moisture would penetrate inside the packages, and

active devices could suffer corrosion damage. This would not only lead to

the destruction of an expensive hybrid, but more importantly, the entire

electronic system might fail.

While lead fatigue is well understood and well characterized, the kinds of

forces that are required to produce hermetic-seal damage in hybrids have never been determined. The flatpack"*" type of package is generally con- sidered to be more susceptible to hermetic-seal damage under various conditions than the plug-in type of package. The vibration experiment included

both types of packages and was designed to determine whether lead fatigue or hermetic seal integrity is the most sensitive to vibration. Hermetic damage

is more difficult to assess than lead fatigue since electrical measurements

of the device and visual observation of the seals are of limited use. Therefore, AE instrumentation was used during vibration experiments to detect

glass cracking in real time, and final determination of seal damage was made

after vibration with a helium leak detector at both room temperature and

~1 25°C.

Before actual vibration tests began, a review of the literature was conducted

in order to avoid repeating any earlier experiments. This was done both in

the NBS library and at the DoD Shock and Vibration Information Center. It was found that most of the relevant work on vibration of electronic components had been done a number of years ago and the investigations were limited

fatigue damage. In only one recent paper was the possibility of cracked

glass seals suggested [14], but without photographs or other supporting evidence. The most convenient and complete single reference to vibration effects on electronic systems is a book by Steinberg [15] which has a chapter

on the vibration of PC boards and other chapters on the fatigue of wire leads

in components. Steinberg has also written a series of articles in trade

magazines [16-21], most of which appear to be derived from his book. Other

relevant papers in the archival literature or presented in conferences [22-

26] were also reviewed. In all of these references, the resonant frequency

* Results and conclusions of this work were presented at the 51st Shock and

Vibration Symposium, San Diego, California, October 21-23, 1980,

+ This is sometimes called a butterfly package.

4
calculations for PC boards were based on simplified equations and were limited to the lowest (the "natural") vibration mode, which is typically in the

100- to 300-Hz range. Since military vibration requirements extend to 2000

Hz, it was necessary to derive more complex equations to obtain correlations

with actual vibration measurements and to be able to locate relatively benign

component mounting positions on PC boards.

A-1.

2

Theoretical Considerations

The vibration modes of a printed circuit board may be obtained if the board

is modeled as a vibrating plate. Various equations for the vibration modes

of plates and membranes having specified boundary conditions have been derived in the literature [27-30]. The most commonly used mathematical methods

were set forth by Rayleigh [31]. The boundary conditions are the dominating

factor in determining modal frequencies; different conditions can result in

frequencies which differ by an order of magnitude.

The three boundary conditions appropriate to PC boards are "fixed," "supported," and "free." For the fixed conditions, sketched in figure 1, the board

edges are firmly clamped such that the edge motion is zero and the tangent is

also zero. The term "supported" (also referred to as "simply supported" or "freely supported") is a carryover from beam-bending theory where the ends of

the beam are supported on knife edges. For a PC board, each edge is considered supported on top and bottom between two knife edges. In this case,

motion at the edge supports is zero but the tangent is not constrained, as shown in figure 2.

Figure 1 . Sketch of "fixed" boundary Figure 2. Sketch of "supported"

conditions. boundary conditions.

The resonant frequencies of fixed boards are significantly higher than those

of supported ones because the fixed mount effectively reduces the length of

board available for vibration. The free case will not be considered here

since all PC boards used in avionics are clamped or supported in some manner.

As stated earlier, previous analysis of PC board vibration was done with

simplified equations which did not permit calculation of higher resonant

modes. In the present work, general equations were derived using, as a basis, the models of Warburton [28] and Hearmon [30]. The dimensions used in

the equations are identified in figure 3.

5
m Number of nodes along dimension a, n = Number of nodes along dimension b,

For fixed boundary conditions,

m = n = 2 for lowest mode.

a h

Figure 3. Sketch identifying dimensions of a vibrating plate.

For the fixed (or clamped) edge board, the frequency equation is

f =

12TrEgh'

12p(l - a^]

+ ^ +

2_jr£ _ 2) (e - 2)

1/2

L a a2b2

where:

(1 )

g = gravitational constant,

E = Young's modulus,

a = Poisson's ratio,

p = board volume density = W/abh,

W = board weight,

Y = (m - 1/2)tt, and

e = (n - 1/2) TT.

For the clamped board, the fundamental mode occurs for m = n = 2 (a node at

each end of the board).

For the supported PC board, the frequency equation is

Egh'

f = 2a2 12p(1 - 0^)

a'^ £^2 2a2 Yi2 e^2

1/2

(2)

where = ~ 1 ^i^<3 e-j^ = n - 1.

The sketches in figure 4 give pictorial examples of the first several vibration modes of boards having various a/b ratios. These were calculated for

various boundary conditions as indicated.

To calculate the vibration modes to be expected in the first experiments

(where boards that were fully copper clad on one side were used), the following values for E and e are composites of the appropriate thicknesses of epoxy

board and copper cladding (all in English units):*

* Dimensions and other values are generally given in English units since

these are used exclusively in this field. Values of E and a for epoxy

boards are ^given in reference [15].

6
(a) The first four vibration modes of a simply supported plate. (From Morse

[27] .

/

?

/

/ o

^\' D /

'

1 '

; ; V /

' / y / / / / y J

' X' \. /

' X. / / X

' / N

; b

' U ;

(b) Vibration patterns of a square plate with fixed edges.

(c) The variation of the vibration pattern as the length-to-width ratio

changes on a plate having free edges.

a/b = 1.00 a/b-0.98 a/b = 0.95 a/b = 0.90 a/b=0.75 a/b = 0.50

(d) The variation of the vibration patterns of a plate having two fixed and

two free edges as the length-to-width ratio changes. These patterns

also simulate the effect of nonuniform loading of a PC board. (From

Warburton [2 8])

Figure 4. Pictorial examples of the first several vibration modes of rectangular boards

7
E =

O = g

=

2.8 X 10^ psi,

1.13, and

385 in./s^.

The dimensions of the board were a = 8,2 in., b = 5.2 in., h = 0,062 in., and

W = 0.21 lbs. The outside dimensions of the boards were typical of those

used in avionics equipment (6 by 9 in.) reduced by the "fixed" clamp supports. The first several modal frequencies for fixed and supported edge

boundary conditions are listed in table 1 . The last column in the table

gives the modal frequencies of a commonly used smaller board of the same thickness.

For comparison, if the constants appropriate to an epoxy board with no copper

cladding (E = 2 x 10^ psi and a = 0,12) were used, the modal frequencies

would all be approximately 16 percent lower. Although some of the frequencies in table 1 may appear to be rather high, these were included because

they may be excited by harmonics of lower frequency vibrations entering avionics equipment, as well as by "white" vibration transmitted from such

sources as jet engines. Also, examination of eqs (1) and (2) shows that the

modal frequencies of a uniformly loaded board decrease as the square root of

the board's weight. Thus, since a board containing from 5 to 1 0 components

(hybrids, capacitors, etc.) can easily weigh 2 to 4 times that of the unloaded board, the frequencies may be as little as 1/2 those calculated in

table 1 . It should be noted that all possible vibrational modes will not

necessarily appear during actual PC board vibration.

Table 1 . Calculated Modal Frequencies for Two PC

Boards

5.2 X 8,2 in. 4x6 in.

Modes (W = 0,21 ) (W = 0. 1 1 8)

m n fixed f

supp fixed

2 2 330 Hz 170 Hz 570 Hz

3 2 495 320 885

2 3 820 540 1400

3 3 975 690 1685

4 2 775 570 1410

4 3 1235 935 2180

4 4 1960 1550 3400

3 4 1710 1300 2930

A-1 .PC Board Layout Considerations

Sketches such as those shown in figure 4 may be used by the PC board layout

engineer to minimize vibrational damage to attached components. Obviously,

it would be desirable to solder components at nodes; however, a node for one vibration mode could be a maximum for another. Higher order modes usually

produce lower vibration amplitudes and, even though the frequency is higher

8
(more excursions per second), result in less mechanical flexural damage to

devices and their leads. Thus, only the four or five lowest order modes need

to be considered with the lowest one being the most significant. Therefore,

component lead placement should avoid straddling the maximum of the lowest

order mode, which is across the shortest dimension of the PC board. Thus,

generally, component placement in the center of the board should be avoided.

It is preferred that the devices straddle a node of higher order modes along

the length of that node (while avoiding the maxima of the lowest mode). One

should especially avoid placement in positions where leads will experience

torsional forces since torsion can more easily damage glass-metal seals than

linear flexing. The designer should also be aware that a board asymmetrically loaded with components (e.g., having the weight concentrated at one end)

will not vibrate in the symmetrical manner shown in figure 4c, but may more nearly approach modes shown in figure 4d. Thus, symmetrical loading is desirable, but if it is not possible, then various stiffeners or snubbers

should be used [20,25,32] as compensation.

A-1 .4 Experimental Measurements

A- 1 .4.1 The Equipment Setup and Preliminary Vibration Tests of PC Boards

The complete experimental setup consisted of a programmable frequency driver;

an Unholtz-Dickie* vibration shaker and control unit Model 106 SP7; two

accelerometers , one for the shaker head and a 1 -gram weight unit to be

attached to the board or hybrid; a double-pen recorder to simultaneously

record both the shaker and the board's acceleration, as well as a special

slip-synchronized strobe light to aid visual observation. A photograph of

the overall experimental setup is given in figure 5.

In consultation with avionics equipment authorities, it was determined that a "fixed" mount for the PC boards best simulated mounts found in the avionics

equipment of interest. An appropriate mount was designed and constructed for use in vibrating boards along two major axes. Design considerations were obtained from the literature [33,34] as well as from consultations with ex- perts who have worked in the vibration field.

Several preliminary runs were made for calibration and equipment setup purposes. The equipment was programmed to follow curves required in MIL-E5400R, such as is shown in figure 6, curve Ilia.

Figure 7 gives the resonant frequency plot of an unloaded copper-clad PC

board with dimensions of 8.2 by 5.2 by 0.062 in. The resonant peaks approximate those calculated for the "fixed" board in table 1.

Frequencies obtained from the chart, along with the appropriate nodes, are shown in table 2.

* Certain commercial equipment, instruments, or materials are identified in

this paper in order to adequately specify the experimental procedure. Such

identification does not imply recommendation or endorsement by the National

Bureau of Standards, nor does it imply that the materials or equipment

identified are necessarily the best available for the purpose.

9
Figure 5. An overview of the vibration equipment. The PC board is shown in

a fixed mount on top of the shaker which is in lower right of figure.

1.0
FREQUENCY (Hz

CURVE la (OF PREVIOUS ISSUES) - REPLACED BY CURVE IVa CURVE lla - EQUIPMENT DESIGNED FOR OPERATION ON ISOLATORS WITH

ISOLATORS REMOVED

CURVE Ilia - EQUIPMENT MOUNTED IN FORWARD HALF OF FUSELAGE OR

IN WING AREA WITH ENGINES AT REAR OF FUSELAGE

CURVE IVa - EQUIPMENT MOUNTED IN REAR HALF OF FUSELAGE OR IN WING AREA WITH WING OR FRONT MOUNTED ENGINES

CURVE Va - EQUIPMENT MOUNTED IN ENGINE COMPARTMENT OR ENGINE PYLON

CURVE Via - EQUIPMENT MOUNTED DIRECTLY ON ENGINE

Figure 6. The sinusoidal vibration requirements for equipment designed for

installation in jet airplanes (from MIL-E-5400R)

.

11
120

110

100

90

g 80

o

0° 70

u

^ 60

^ 50

40

30

o 20

10

0

/

10

"1

I I I I r~i

[

"1—I—"1 r

PEN DISPLACEMENT

I I I I I I I J I I I 'III

4 9=

3 S

-2

1 ^

100 1000

FREQUENCY (Hz)

Figure 7. The resonant vibration peaks on a 20-min frequency sweep of an unloaded PC board (curve a). The recorder pen was displaced '^^l/S in. to the

left to prevent intercepting the calibrating pen that was recording the

shaker drive (curve b) . Saturation on the resonances is evident on three

peaks

12
Although the shaker drive was constant at 5 g, the accelerometer on the board

recorded troughs of 10 g, which presumably resulted from overlapping resonances. Nevertheless, the board has an amplification factor ( transmissibi lity) of >10 at some resonances. It is apparent from table 2 that there is

excellent agreement with theory.

Table 2. Comparison Between Calculated and Measured Modal Frequencies for an Unloaded Board that was Copper Clad on One Side.

m n

f meas ^calc Comments

2 2 310 Hz 330 Hz Peak > 100 g

3 2 480 495 Very small peak

2 3 800 820 Peak >100 g

3 3 960 975 Peak ~ 80 g

5 2 1020 1 160 Peak > 1 00 g

5 3 1640 1603 Peak near 100 g

3 4 1 760 1710 Peak > 100 g

A-1 .

Preliminary Experiments on Vibration of PC Board-Mounted Hybrid

Packages

In order to prepare flatpack hybrid packages for mounting on boards, it was necessary to bend and cut the leads. Package manufacturers recommend bending

leads no closer than 0.10 in. from the package and they prefer a 0.15-in.

spacing [35]. Subsequently, two lead bending-cutting jigs were designed and

constructed. One bent the leads 0.07 in. and the other 0.15 in. from the

edge of the package.

For this particular experiment, eight packages had their leads bent and cut,

four on each jig. All showed a small degree of visually observable glass

damage to two or three out of 30 leads, but this was much less than on prebent samples supplied from outside. Two high-quality packages, one with

0.07-in. and one with 0.15-in. leads, were soldered to unetched PC boards.

This was facilitated by scribing a moat with a sharp knife in the copper

cladding around the area to be soldered. The thermal conductivity away from

the leads was thereby decreased, and soldering with a small iron was possible. The open packages were mounted side by side in the center of the board

with the leads in the short direction of the board. This was intended to be

the worse position for damage by excitation at the lowest mode. An accelerometer which weighs one gram was attached to the package with the shorter

lead spacing.

The "fixed" nature of the board mount was verified by observation under slipsynchronized strobe light, as well as by observation of lycopodium powder

patterns on the board after a preliminary run. Resonant amplitude acceleration peaks were measured both with the accelerometer and by the peak displacement method using a depth-gage micrometer mounted on a fixed nonresonant

portion of the mount. The lowest resonant mode of the board with two

13
soldered-in hybrids was 280 Hz, which compares well with the calculated value

of 308 Hz (including the weight of the packages, solder, and accelerometer )

.

Had this weight been more evenly distributed, presumably the two frequencies

would have been even closer.

During the test, several frequency sweeps were made from 1 0 Hz to 2 kHz;

however, the programmed rate of frequency change was varied to allow longer

observation with the strobe at the resonances and to make amplitude measurements. At several resonances, clear differential motion was observed between

the packages, with the greatest motion occurring on the one having the longer

leads. At one point, package rotation was observed, and at higher modes

differential motion from one side of the package to the other was observed.

After the third frequency sweep, different vibration noises were heard. It was concluded that some lead damage had been sustained, and the test was terminated. The damage was extensive both to the leads, by fatigue, and to

the glass seals as observed under a microscope. Eighty percent of the leads

on the package with longer lead spacing were damaged or broken, as is evident

in the photomicrograph in figure 8. No leads were broken on the package with

shorter leads, but some glass cracks were observed.*

The packages were then removed from the board for further observation of the

damage and evaluated by the newly developed high-temperature leak test method, which is described in Part B of this report. The room temperature leak

test showed both packages to have leak sizes less than 10"^ atm»cm^/s helium.

Tests at 125°C gave similar results, leak sizes less than 5 x 10~^ atin*cm^/s

helium, the test limit. The packages were put on a thermoelectric temperature cycling system and cycled 25 times from -55° to 125°C with a 7-min cycle

time. They were retested, and again the package with shorter leads had a leak size less than lO"^ atm»cm^/s helium at room temperature and 5 x 10~^

atm»cm^/s helium at 125°C. However, the package with the longer leads, which

had sustained considerable damage, showed a 10"^ atm»cm^/s helium leak at one seal at room temperature and 10"^ atm*cm^/s helium leaks at two seals (one was the room temperature leak) at 125°C. It should be noted, however, that a 10-^ leak is acceptable in Method 1014.2 of MIL-STD-883 [36] for packages of

volume 2 cm ^ or greater. Thus, these packages cannot be considered to have

failed hermetic tests even though 1/2 of the leads had been fatigued off.

The above preliminary work was performed on copper-clad boards to establish

techniques, vibration amplitudes, and the accelerations required to induce

visual glass seal damage and lead fatigue.

A-1 .6 The Vibration Characteristics of Metallization-Patterned PC Boards

with Soldered-In Hybrid Packages

When pattern-etched PC boards were available, a series of experiments were

run on them in both loaded and unloaded condition with accelerometers attached to the boards and to the hybrids at various board positions. A typical board with three flatpack hybrids soldered into it is shown in figure 9.

* All further flatpack vibration tests in this study were carried out with

the shorter {0.07-in.) bent leads which gave better fatigue resistance than

the longer leads.

14
(b)

Figure 8. Photomicrographs of a 1-in. square hybrid package after vibration

testing. (a) Leads fatigued off during vibration tests at the package seals,

showing glass damage. (b) Leads fatigued off at the bends or at the solder

joints

15
Figure 9. A closeup of the vibration mount and the printed circuit board with

three flatpack hybrids soldered on it. The accelerometer that weighs only- one gram is on the upper hybrid, and a TO-lOO acoustic-emission detector (sec.

A- 2.1) is shown on the lower hybrid.

16
Figure 10 shows the acceleration versus frequency for two frequency sweeps

made on an etched PC board that contained no^ hybrids and was only loaded with

a 1-gram accelerometer placed in two different positions. The solid curve

was obtained with the accelerometer in the center of the board on the position of the middle hybrid in the previous figure. The dotted curve is taken

with the accelerometer placed on the side position of the board. The resonance peaks have quite different amplitudes depending on sensor placement.

This is to be expected since the vibration amplitude varies considerably from

point to point on the board when various vibration modes are excited. This

particular board was driven with only 1 . 5-g constant acceleration. The

transmissibility in the center of this board for the first or fundamental

vibration mode is almost 100. It is even higher for some other modes. Differences between the various modal frequencies in figure 10 and those calculated in table 1 are attributed to changes in E, a, and W resulting from

patterning the copper film and introducing large numbers of pla ted-through

solder-covered vias.

Figure 11 shows the vibration amplitudes (given as acceleration) as a function of frequency for a printed circuit board with one 15-gram plug-in-type

hybrid package mounted in the center of the board. The solid curve represents data obtained with the accelerometer mounted in the hybrid. The dashed

curve is a result of placing the accelerometer in an empty side position on

the board. It is to be noted that in this position the unloaded board (fig.

10) showed a higher acceleration for the center placement; however, in this case the side placement results in a higher acceleration due to the loading

in the center. An even greater change is observed in the second vibration

mode in which the side accelerometer position shows a much higher acceleration than the center position.

Three plug-in- type hybrids were placed on the board and vibrated. The data

are shown in figure 12. The solid curve represents the acceleration measured

by placement of the accelerometer in the center hybrid position and the highest peak is the fundamental. The second solid curve peak is about half the

amplitude of the first and successively higher frequency peaks are of lower

amplitude. Thus, as is apparent from eq (1), the placement of several hybrids on the board loads the board and lowers the frequency and amplitude of

each modal peak. The constant acceleration drive for this experiment was 3 g, twice as high as that of the unloaded board in figure 10. The dashed

curve represents the placement of the accelerometer on one of the side hybrids. The situation is quite different from the center position, the fundamental is lower, and the second resonant peak is as high in amplitude as that

measured in the solid curve.

A-1 .

Susceptibility of Plug-in-Type Packages to Fatigue Damage at High-g

Levels

Preliminary experiments showed that all f latpack-type packages regardless of

lead length would suffer some fatigue breakage of leads when subjected to

resonant vibration in the order of 1 00 g for a period of 30 min. Therefore,

it was decided to perform similar evaluations of the plug-in-type packages.

The same PC board packages vibrated in figure 12 were further stressed at the

1 00-g level (center hybrid) at tVie fundamental resonant frequency (~250 Hz)

for 1 hour. There was no apparent damage and so the experiment was continued

17
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at 450 g for 30 min longer. Observation with a strobe revealed that the

hybrid packages acted as board stiffeners. The board moved only vertically

(up-and-down) in the vicinity of the hybrids with no flexing occurring under

them.

After these high-stress experiments, the hybrids were removed from the board.

No visual seal damage was observed, and leak tests revealed no loss of hermeticity as a result of this extreme stress test. An additional stress test

was performed at 450 g for 30 min with two plug-in-type hybrids on the board,

one in the center and the other on one side to assess the results of asymmetrical board loading and possible torsional forces on the leads. Again, no damage occurred to the hybrids. However, the 1 -gram accelerometer (rated for

1000 g) that was used to monitor the experiment was damaged due to its leads

fatiguing and had to be repaired before other experiments could proceed.

Even though soldered-in, plug-in packages appear to be immune from vibrationinduced damage, there is still a possibility that seal damage can result from

thermal shock during soldering or desoldering to remove them from the board.

For experimental purposes, open hybrids (as in fig. 10) should be leak tested

on the board before and after vibration.

A-1 .Long-Term Low-g Nonresonant Vibration Effects

Since low-g nonresonant vibration may occur in avionics environments, one test was run to assess fatigue and/or glass seal damage under low stress,

extended time conditions. Four flatpack hybrid packages with short (0.070

in.) bent leads were soldered onto a prepared (etched) board (three symmetrically on top and one in the center on the bottom) and vibrated at 90 Hz for

20 h at the 10- to 1 5-g level. They were removed, visually inspected, and

leak tested. No glass cracking was observed and no leaks at >10~^ atm«cm^/s

were found.

A-1 .Effect of Very High-g Forces on Flatpacks

Experiments were performed to determine whether seal damage in high-quality

packages could be induced when the leads were fatigued off rapidly and violently at high-g levels. Two flatpack packages were soldered onto a PC board

symmetrically, near the center, and vibrated at 200 to 300 g for several

minutes until all of the leads of one package fatigued off and the package

fell from the board. The other package remained on the board held by three

or four (out of 30) leads. Figures 13a and 13b are scanning electron microscope (SEM) micrographs of two of the severely damaged but still hermetic

seals 1 Figure 13c is included from an earlier vibration run at a lower

stress level (~80 g for 10 min). All of the packages were tested for leaks

at room temperature and at 125°C and found to be better than 10"^ atm»cm^/s.

An important conclusion of this experiment was that no correlation was observed between the visual appearance of the seal and leaks when high-quality

packages were studied [37].

21
(c)

Figure 13. SEM photomicrographs of glass-metal seals in hybrid packages that

have been vibrated. No hermetic damage occurred to the packages. (a) and (b)

Seals from packages vibrated at 200 and 300 g at 250 Hz. (c) Seal and lead

from a package vibrated at 80 g at 260 Hz. The small dots on the glass portion

are foreign particles and not a part of the seal.

22
A-2 Acoustic-Emission Equipment and Techniques for Detecting Glass Seal

Cracking During Vibration

A-2 .

1

Development of an Acoustic-Emission Detector for Use in Vibration

Experiments

When acoustic-emission tests are used to establish the after-production integrity of microelectronic welds and seals, the sensor may be attached to a printed circuit board (thickness ~0.06 in.) or to an electronic package such

as the casing of a hybrid microcircuit which may have wall thicknesses of

only ~0.015 in. The AE wave detected in a test generally has a wavelength on

the order of 0.4 in. and propagates along these thin materials as some form

of surface wave such as Rayleigh and Lamb waves. Normal acoustic-emission

detectors are relatively insensitive to these surface waves.

AE detectors are usually constructed with the piezoelectric transducer (PZT)

element recessed into the case and with plastic extending to the surface for

acoustic wave coupling. These sensors are typically designed to have maximum

sensitivity in the direction of the thickness of the PZT disc. Thus, maximum

sensitivity occurs for waves entering the detector normal to its surface

rather than for a surface wave propagating along the base or top of a hybrid

package. Acoustical wedges have been used to increase the sensitivity of AE

detectors to surface waves. However, a disc-shaped PZT sensor lying flat on

the surface is sensitive to surface waves propagating along its diameter,

and the resonant frequency of the sensor to such waves is determined by the

diameter rather than the thickness of the sensor. Also, the impedance of AE sensors is relatively high (~20 KQ, at resonance and higher off resonance),

and the output for weak signals is low, in the order of microvolts. Some of

the signal may be lost due to cable capacitance; in addition, shielded cable

noise generated during PC board vibration tests may limit the maximum sensitivity of the detectors. In order to complete objectives of this study, it was necessary to develop an improved sensor to overcome the above limitations .

A detector-preamplifier having high sensitivity to surface waves was con- structed as follows: AE sensors with diameters of about 0.25 in. were ultrasonically cut from larger diameter 1-MHz and 5-MHz PZT ceramic stock. Copper

wires, PTFE-coated and 0.003 in. in diameter, were soldered to the silvercoated faces of the ceramic disc, and the element was bonded into the cap of

a TO-100 IC package, which had a thin layer of polyimide in it as insulation.

A 733 differential input-output operational amplifier chip was attached with

insulating epoxy to the base of the TO-100 package and wire bonded so that it

had a gain of 100. The chip was cleaned with fluorocarbon and DI water and

then protected with a film of xylene-thinned silicone rubber. The bonding

wires were protected from possible shorting by a film of photoresist. Photographs of the basic assembly steps are given in figure 14.

The completed AE detector-amplifiers weigh approximately 1.3 grams and,

therefore, offer negligible damping when attached to hybrids undergoing vibration tests. All connections to the device, including signal leads, are unshielded twisted pairs in which vibration-induced noise is negligible at

the relatively low output impedance of the 733 chip.

23
Figure 14. Steps in the assembly of the TO-lOO acoustic-emission detectorpreamplifier. From left to right, the ijA-7 33 chip is die bonded and wire

bonded; then a silicone rubber coating material is applied. Next, a PZT AE

detector chip is shown. On the right it has leads attached and is held to

the TO-lOO cap with a polyimide adhesive. The leads are then ultrasonically

attached to the package, the cap is welded on, and external power and signal

leads are attached. (A closeup of a finished detector is shown in fig. 21.)

24
The amplitude versus frequency response for surface waves of several of these

acoustic-emission detectors was determined by calibration at NBS.* Figure

15 is the amplitude versus frequency curve for one of the detector units that

contains a 1 -MHz PZT (in the thickness mode). Note that there are high re- sponse peaks at approximately 0.2, 0.4, and 0.7 MHz with only small peaks oc- curring around 1 MHz, the thickness mode response peak. The method of mounting enhances the response to surface waves traveling parallel to the diameter .

Figure 16 shows the amplitude versus frequency response of a detector whose

PZT has a 5-MHz peak response when used in the normal or thickness mode. The

diameter response mode is around 0.4 MHz. The amplitude at this frequency is

considerably smaller than that of the first detector, but there are three

clear peaks in the 0.35- to 0.4-MHz range. No other response of any significance is evident in the indicated frequency range. For comparison, figure 17

shows the frequency response of a high-quality acoustic-emission detector

obtained from a commercial source. This has a peak at approximately 0.3 MHz.

It was chosen to peak in the same frequency range as the other two detectors

as well as having a high sensitivity as stated in the manufacturer's specifications. The commercial unit which contained no built-in 40-dB preamplifier

was plotted on a different relative intensity scale from the two NBS detectors. The three curves (figs. 15, 16, and 17) are made equivalent if one multiplies the commercial detector response (fig. 17) by 1.35. This makes

its peak magnitude approximately 3 on the scale of figures 15 and 16. The

first detector is clearly far more sensitive in the 0.3-MHz range (in the

order of six times) than the commercial unit. The second TO-100 detector has

a peak sensitivity somewhat less than the commercial unit, but the response

is much broader, in the 0.3- to 0.4-MHz range. Therefore, that TO-100 detector should be more sensitive to a typical acoustic-emission burst.

A-2.2 Determining AE Waveform Signatures in Order to Separate Glass Cracking

From Metal Fatigue

The TO-100 detector whose calibration was given in figure 15 was chosen to

determine the waveform signature of glass cracking. This detector was acous- tically coupled to a hybrid flatpack containing 30 leads. To obtain an acoustic-emission signal of glass seal cracking, one of the leads was intentionally bent upward until the glass cracked. The acoustic-emission signal

was fed into a transient recorder. This recorder was interfaced to a computer, which performed a fast Fourier transform (FFT) on the data. The data were analyzed in three separate parts: (a) the first third of the waveform

that included the leading edge of the AE burst, (b) the middle portion, and

finally (c) the remaining third. The wave form analysis was divided to determine any differences in the frequency output as a function of time. Figure 18a is the FFT of the leading edge of one of these glass-seal-crack

acoustic-emission bursts. There are two high peaks, one near 0.2 MHz and one at about 0.4 MHz. There is also a moderate signal in the 0.6- to 0.7-MHz

region. The amplitude of this latter signal was not corrected for the sensitivity characteristics of the detector. If it had been, the 0.7-MHz signal

would be much larger.

* The calibrations were carried out in the Mechanical Production Metrology

Division of NBS by F. R. Breckenridge

25
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Figure 15. The frequency spectral response of a TO-lOO acoustic-emission

detector-preamplifier. (The PZT ceramic element used in this test was designed for a 1-MHz peak response.)

26
T

Figure 16. The frequency spectral response of a TO-lOO acoustic-emission

detector-preamplifier. (The PZT ceramic element used in this test was designed for a 5-MHz peak response.)

27
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compared

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28
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Figure 18. Fast Fourier transform (FFT) of an acoustic-emission waveform obtained by intentionally bending a lead on a hybrid until the glass seal cracked.

The FFT was divided into three equal parts [ (a) leading edge, (b) center, (c)

trailing edge] to show any differences across the entire AE waveform. The ar- row in (a) indicates the 700-kHz peaks that vanish after the first third of the

AE burst.

29
Figure 18b shows the FFT of the middle third of the acoustic-emission burst.

As with most AE bursts, the amplitude decreases with time across the burst.

Therefore, the overall amplitude of this signal is somewhat lower than the

first. One should note, however, that even though the amplitude is lower,

the relative amplitude of the various peaks is significantly different. In

this case the 0.2-MHz peak is relatively lower than the 0.4-MHz peak, and the

0.7-MHz peak has almost vanished. Figure 18c shows the last 1/3 of the FFT

frequency spectrum. In this case, the overall amplitude is lower and there

is _no signal in the 0.7-MHz range. Also the relative amplitude has again

changed between the two major peaks. They are now approximately equal.

Similar data to these were obtained with other glass cracks, and although the

exact shape of the FFT peaks was different in each case, a shift to lower

frequencies was observed from the leading edge to the center of the waveform.

It had also previously been determined that the package metal-fatigue

acoustic-emission signals did not produce the same high-frequency component

in the leading edge, and in addition the overall amplitude was lower with

more continuous AE rather than high bursts. Therefore, the triggering

circuit that is used to activate the transient recorder was preceded by a filter that only passed the 0.7-MHz signal. Thus, by using the 0.7-MHz

filter, it was possible to separate AE originating from lead fatigue from

that produced by glass cracking. This separation had originally been thought

to be a major problem since both fatigue and cracking produce acoustic

emission.

A-3 Acoustic-Emission-Monitored Vibration Experiment

A-3 .

1

Equipment Setup

The first step in the AE-monitored vibration experiment was to establish a procedure to attach the TO- 100 detectors to the hybrids. For this, it was necessary to determine the optimum material that would serve both as an AE

couplant as well as a mechanical glue-like attachment for use in holding the

detectors during vibration tests. Several thermoplastic materials typically

used as AE couplants were rejected because of their brittleness, since any

cracks that developed in the couplant due to the vibration could generate

acoustic emission and this might be confused with signals emitted from cracking seals. A series of tests involving a dozen different materials determined that beeswax would adequately serve as both a mechanical and an AE

couplant for vibrational purposes. The AE detector can be attached either by

melting the beeswax (~70°C) or by solvent evaporation (the former is most

desirable since it is quicker).

The AE equipment setup that was used to control and record the hybrid sealcracking vibration experiments is given in figure 19. The output from the

acoustic-emission detector can completely control the experiment. As seen in

the figure, the shaker drive oscillator output was fed through an electronic

trigger-controlled switch. This switch was triggered to the of f -condition by

a pulse from the acoustic-emission trigger unit. Thus, when an acousticemission burst was obtained, the sine wave drive to the shaker was removed

within a few microseconds. Simultaneously, the acoustic-emission burst that

initiated the trigger signal was recorded by the transient recorder, and the

accelerometer output, which was plotted on an x-y recorder, displayed the

30
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31
acceleration that the package was undergoing. The AE burst was later photographed from the oscilloscope monitor and a complete record of the experiment

was obtained.

The vibration equipment was described earlier and shown in figure 5, and

figure 9 was a closeup of a PC board in the "fixed" vibration mount. Three

flatpack hybrids were shown soldered on the printed circuit board. Figure

20 shows a comparable situation in which three plug-in-type packages are mounted on a printed circuit board. The acoustic-emission detector is attached to the hybrid on the left. The accelerometer is attached to an empty

hybrid package in the center. An additional hybrid is on the right of the

board. A closeup of the TO- 100 detector mounted in a hybrid is given in

figure 21 . The general acceleration versus frequency data for this board

configuration were given in figure 12.

New (unvibrated) f latpack-type hybrids were soldered into patterned printed

circuit boards in order to assess seal damage in real time. Some of these

hybrids had intentionally damaged glass seals that would have failed a visual

inspection, although they passed leak tests at both low and high temperatures. Other hybrids that had almost perfect glass metal seals, both visually and in leak tests, were also mounted in order to assess correlation between initial visual characteristics and any hermetic damage resulting from

the vibration.

Preliminary work (sec. A-1.7) had shown that high-quality plug-in-type packages that were soldered into PC boards remained undamaged (either by lead

fatigue or hermetic leaks) when vibrated at high accelerations. Therefore,

it was decided that all acoustic-emission experiments would be carried out on

f latpack-type packages that were more subject to damage by vibration. The

packages were split into two groups. In one group, there was essentially no glass cracking visually evident. The other had a significant amount of glass

damage intentionally induced by moving the leads back and forth until large

pieces of glass had cracked from the outside of the seal area. The packages

were verified as being hermetic before soldering onto PC boards. Vibration

tests were run separately for each package so that none would undergo acceleration when it was not being monitored by AE detectors.

Figure 22 is a composite of the x-y recorder curve tracing showing the acceleration experienced by an initially good package (no glass cracks by visual

inspection) as it was scanned through its frequency spectrum until a glass

crack occurred. The photographic insert shows the acoustic-emission burst

that occurred at the peak acceleration. The acoustic-emission trigger re- moved the shaker drive, and the acoustic-emission burst was simultaneously

recorded by the transient recorder (see fig. 19). After this first glass

cracking occurred in the essentially perfect package, further runs produced

AE bursts at lower accelerations, in the order of 1/3 to 1/2 that of the

first run. Equivalent tests were also made with intentionally damaged packages. In no case during these experiments did the acceleration ever reach

levels higher than 40 to 50 g before acoustic-emission bursts occurred.

After the AE tests, all of the packages were tested on a helium leak tester,

and even though glass cracking had occurred, there were no hermetic leakers.

32
Figure 20. A printed circuit board with three plug-in-type hybrids soldered

to it. The hybrid at the left has a TO-lOO AE detector on it, and the center

package has a small accelerometer on it. The board is in the "fixed" mount

which is attached to the vibration shaker.

33
Figure 21. An enlargement of the leftmost hybrid in figure 20 showing more

details of the TO-lOO AE detector.

34
FREQUENCY (Hzj

Figure 22. Composite of the acceleration experienced by a flatpack hybrid

and the acoustic-emission burst emitted at the peak acceleration of the sol curve. The instrumentation turned off the vibration shaker drive when the

burst was received, so the acceleration dropped to zero at that time.

35
In the above test using high-quality packages, it was found that glass cracking could occur under acceleration-type stresses. However, the objective of

this study was to determine whether acceleration was capable of damaging the

hermetic seals on PC-board-mounted hybrids. The normal way to determine

whether such damage actually caused the seals to leak was to perform a helium-leak test. Such leak tests were run on all packages before and after

vibration tests. Although some fatigue-damaged seals were observed to degrade slightly at 125°C after 25 thermal cycles from -55°C to 125°C, they

would have still passed MIL-STD-883 hermeticity tests [36] for that size

package. The above packages all came from lots that were considered excellent, and some selected packages from these lots had been shock tested from

350°C to liquid nitrogen temperature without loss of hermeticity.

The question still remained whether packages from reject lots that were screened as acceptable could undergo vibration and still maintain hermeticity. Lots rejected because of serious hermeticity problems are hard to obtain

since the customer usually returns them to the manufacturer. However, sam- ples from one such lot* called out in a GIDEP Alert ( F3-A-80-01A) [38] were obtained and rescreened for leaks. Two of these packages from the lot were found to be hermetic. These were mounted on a PC board and vibrated for 1 h

at 70 to 80 g. Some fatigue damage was observed and one or two leads fell

off, but neither the lead damage nor visually evident glass seal cracking was

in any way as severe as was observed on some of the high-quality packages

described above. One of the two packages from the reject lot that were initially tested and found to be hermetic became a "gross" leaker after vibration. The sample size was small, but considering that no high-quality package became a leaker during vibration tests, it must be concluded that

screened "good" packages from an otherwise "bad" lot appear to be more sus- ceptible to vibration damage than do packages from initially "good" lots.

A-3 .

2

Conclusions of the Vibration Experiments

Hybrid packages that were soldered into PC boards were vibrated over a range

of frequencies and accelerations. High-quality plug-in-type packages are essentially immune from vibration damage and appear to act as board stiffeners. Several of these packages sustained 450 g at 250 Hz for 30 min with no lead fatigue and no glass cracks or hermetic-seal damage. Normal conductors

on PC boards would begin to fail at these accelerations. Therefore, highquality plug-in-type packages soldered directly into plated-through holes in

PC boards should be used where practical in equipment that may be subject to

vibration.

The leads on f latpack-type packages are subject to vibration fatigue damage.

The possibility of damage increases as lead lengths increase. However, the

process of bending leads close to the glass seal may itself produce meniscus

cracks. Care should be taken to design lead-bending jigs carefully in order

to minimize this problem. If long leads are required, they should be protected by epoxy-bonding the cases to the board or by using appropriate board

snubbers to dampen vibration.

* These flatpack samples were obtained from J. R. Beall, Martin Marietta

Corporation, Denver, Colorado.

36
During flatpack vibration experiments, glass cracking (meniscus and other)

starts at about 100 g for good packages and as low as 20 g on good packages

with intentionally damaged glass seals, as revealed by real-time acousticemission measurements. When tests were run on good packages, there were no cases where the glass damage penetrated through the depth of the seal and

caused hermetic leaks, even when leads were fatigued off rapidly and violently at 200- to 300-g levels.

Hermetic damage resulting from vibration was observed on one package out of

two from a reject lot that had passed a hermetic test screen. This suggests

that screening and then using the "passed" packages from a lot that had a high percentage of failures (a "bad lot") may lead to reliability problems.

No lead fatigue or glass-seal damage has been observed on high-quality PCboard-mounted f latpack-type packages from either resonant or nonresonant

vibration at 10- to 1 5-g levels for extended periods (approximately 20 h).

A major conclusion of this work is that no correlation has been observed

between visually obvious glass-seal damage in hybrid packages and hermetic

leaks, as measured by a helium-leak detector, at room temperature and at

1 25°C.

37
PART B NONDESTRUCTIVE STRESS TESTS FOR HYBRID PACKAGE INTEGRITY

B-1 Leak Test Measurements and Procedures*

B-1 .

1

Development of a High-Temperature Open-Package Leak Test Procedure

It has been reported by Thomas [39] that some packages may leak at high temperatures or during rapidly changing temperatures, such as encountered in

temperature-cycling tests, but are hermetic during normal room temperature

leak tests. There are two reasons for this. The first is that the kovarglass seal is in tension above ~125°C, but in compression at room temperature [40]. Thus, if there are defects in the seals, they may open at high

temperature and result in temporary leaks, and then reseal at lower temperatures where leak tests are usually made. The second reason is that, during

rapid temperature rise, nonequi librium thermal conditions exist with the

glass temperature lagging that of the better conducting kovar, putting any

weak areas of the seal in significant tension, thus causing defects in the

seal. Therefore, in order to fully evaluate seal damage and correlate this

with acoustic-emission bursts, it was necessary to develop a high-temperature

open-package leak test.

The normal open-package leak test, using neoprene rubber and a standard vacuum grease, is not useful at 1 25°C because the rubber can outgas and in addition becomes highly permeable to helium. The grease becomes mobile and moves out from under the package edges.

The first efforts to develop a high-temperature leak test were directed

towards finding published data on the permeation of helium through various

elastomers and greases at temperatures greater than 100°C. No such information was found. The next step was to obtain several different greases and

very thin sections of different elastomers and greases, then leak testing

known good packages and dummies at various temperatures, choosing the best

combination. The present system uses thin Viton A elastomer and Fomblin

grease. This combination is adequate for leak testing open hybrid packages

to 1 0 atm'cm /s up to about 125°C. In addition, a temperature-controlled

package heater was designed and constructed and is shown in figure 23,

Although a number of packages have been successfully tested by this method, new procedures and materials should be investigated in order to make utilization of the test more rapid and practical for monitoring large numbers of

samples

.

B-1 .

2

Preconditioning Procedure for Hybrid Packages Before Hermetic Testing

When leak testing, it is essential that packages be carefully prepared in

order to remove any contaminants that can plug or otherwise clog possible

leakage paths.

Tv/o major interferences in the leakage testing of packages and package components are surface sorption and clogging. All device configurations consisting of glass, metal, and ceramic, or combinations thereof, including coatings

* The work in this section was done by Stanley Ruthberg.

38
Figure 23. The high-temperature open package leak test apparatus. (a) Fluorocarbon elastomer gasket. (b) Platen used to conduct heat uniformly to the

package. A thermocouple controls the temperature on the platen. (c) Heater

element. (d) Hybrid package. (e) Leak detector port.

39
and external sealants, are subject to the adherence of tracer gases by adsorption on the surface and/or by absorption into the material. This effect

can cause a high background indication which leads to reduced test sensitivity or in many instances can be great enough to cause spurious rejection as a leaking component. Fingerprint oils and other contaminants introduced by

careless handling will cause the same effect. Thus, preliminary measurements

should be made on representative specimens, and the dwell time between the

pressurization of sealed packages and detection of helium or other tracer

gasses should be lengthened until desorption has reduced the signal relative

to the test reject level. A gas flush of the package with nitrogen or clean,

dry air can accelerate desorption in some cases.

Very small holes that may give rise to appreciable leakage rates can be

blocked by small solid particles or by condensed liquids. Test results may

be negative; yet, such channels can reopen at a later time because the particles become dislodged or the liquid evaporated. Until leak tests have been

made, the components should be kept out of contact with ambient or other

vapors as well as particulates that could cause blockage. Test samples

should not be held with bare hands !

In general, the retention of a condensed liquid within a microchannel is

governed by surface tension. The differential pressure required to dislodge

the liquid is equal to 2a/r, where a is the surface tension and r is the

channel radius. In the case of an air-water interface, a is ~73 dyne/cm at

20°C and ~59 at 100°C, while the vapor pressure ranges from ~1 7 Torr at 20°C

to one atm at 100°C. As an example, a meniscus in a channel of l-ym radius

requires ~1.4 atm pressure differential to dislodge the water at 20°C and

~1 . 1 atm pressure differential at 100°C. A channel of this radius and 1 -ym

length would give rise to a standard leak rate of ~2 x 10"^ atm«cm^/s. Obviously, the 1 -atm pressure differential normally encountered in a helium leak

test of hybrid packages* is insufficient to dislodge the water in this

instance. At 100°C, the 1 -atm vapor pressure of water produces zero pressure

differential within the channel when the component is in normal ambient, and

the water could only be dissipated by the liquid diffusing out of the channel. This diffusion would occur at a rate approximately two orders of magnitude less than air would pass through this same leak size, considering the

pressure increase inside the package due to gas expansion (~0.2 atm). However, were the component also placed in vacuum, the water would be eliminated

at a rate faster than the standard leak rate since the internal pressure

would add to the vacuum giving ~1.2 atm of driving force. In actuality,

leaks are not uniform channels, but rather they are cracks and torturous

passages much longer than in the example and require greater forces to dislodge any liquid.

Where surface contamination is suspected or where exposure to vapors and

liquids has occurred, a preliminary solvent wash followed by a vacuum bake

(filtered house vacuum is adequate) of two or more hours at 1 00°C or greater

is necessary for both open or sealed package leak tests. Solvents should

have relatively low surface tension and high vapor pressure. For example.

* MIL-STD-883 Method 1014.2 [36] specifies 30 ± 2 psig helium bombing for 4 h

for most hybrid package sizes. The package is normally sealed at 1 atm

(14.7 psig); thus the test is conducted at a 1 -atm differential.

40
Freon TF has about 1/4 the surface tension of water at 20°C and ~4 atm of

vapor pressure at 100°C, These characteristics provide a more appropriate

driving force and dissipation rate under vacuum bake. Even with all of the

above precautions, it is best that components be kept clean and that fine

leak testing be carried out before any liquid contact. However, a bake at

150°C at normal atmospheric pressure or at 1 00°C in a vacuum is a satisfactory preparation for leak testing in order to remove water or fluorocarbon

plugging.

B-2 Development of an AE-Monitored Thermal Shock Test

for Open Hybrid Packages

B-2.Introduction

The published thermal expansion curves of kovar-glass seals [40] indicate

that the outside perimeter of a hybrid package seal is in tension at ~150°C

in the radial direction and either in compression or unstrained at room temperature. The center lead interface would be in compression at 150°C and in

tension below room temperature.

The actual stresses in hybrid flatpack seals have never been characterized.

The nearest configuration studied is the cylindrical seal [41]. The circular

hole and rectangular lead configuration of the hybrid seals produces unknown

stresses during temperature cycling and shock tests. Although such stresses

could be calculated by finite element analysis, this has not been done. In

addition, classical curves comparing the expansion coefficients of kovar

(iron-nickel-cobalt alloy) and typical sealing glasses represent average

rather than specific lot characteristics. Manufacturers' specifications for

sealing glass permit expansion coefficient variations of ±1.5 x 10"^ m/m°C

(from an average value of 5 x 10~^ m/m°C), and for kovar the ASTM specification (ASTM F 15-78) [42] allows ±3 x 10"'' m/m°C variations in the 30 to 400°C

range (from an average value of 4.9 x 10"^ m/m°C). Worst case matching between the glass and metal can produce strained seals that may not tolerate

temperature shock. Hybrid package manufacturers seldom ascertain the exact

expansion coefficients of their raw materials, and thus, even if the kovar

oxidation process is correctly controlled, a certain amount of luck is necessary in order to produce reliable seals. Furthermore, fillers and hardening

additives are often mixed with the glass, and although these are generally

considered to be benign, they have an unknown effect on the expansion coefficient as well as the "wetting" of the kovar oxide. To assure reliable

glass-metal seals, it is essential that package manufacturers evaluate the

residual strain of their specific kovar-glass combination. Methods such as that of ASTM Standard F 14-76 [42] offer simple, inexpensive means of obtaining such evaluation. It is fortunate, and perhaps essential to hybrid reliability, that the dimensions of the package seals are small, limiting the

seal area and therefore the possibility of flaws.*

Considering the above facts, it is apparent that all hybrid package production lots should be screened by a sensitive, reliable test that will reveal

reveal weak glass metal seals. McCormick and Zakraysek [43] described the

effect of kovar oxide thickness on the reliability of glass-metal seals used

* Contributed by H. E. Hagy, Corning Glass Works, Corning, New York.

41
in hybrid packages. They concluded that a shock test (MIL-STD-883 Method

1011.2, Condition C) [36] was an adequate test for hybrid packages. However,

inquiry revealed that many manufacturers routinely use only a hot plate shock

test unless the MIL-STD-883 test is specifically called. Briefly, the hot

plate test is performed as follows: a kovar flatpack containing glass-tometal seals is placed (face up) on a hot plate that is heated to a temperature in the range of 350°C to 400°C. The package temperature is allowed to

reach equilibrium. Following removal and cooling, the glass seals are examined and leak tested for damage. Because of the frequent use of the shock

test as a screen, it would be desirable to instrument this test with AEdetection equipment to obtain rapid information on the seal integrity that

may not be revealed in a visual inspection or even in a room temperature leak

test. Such an AE-monitored test could save the time required for a leak test

and may give other information related to reliability of the package.

B-2 .

2

Characterization of the Hot Stage Shock Test

An acoustic-emission-monitored hot plate shock test would require that the

packages be put on the hot plate in a face-down configuration with the AE

detector on top (attached to the normal bottom of the package). This facedown configuration is thermally different from the usual face-up hot plate

test, so it was necessary to measure the thermal characteristics carefully

for both situations. During these measurements, it became evident that this

frequently used test is not well characterized and under various conditions

could pass marginal packages because the temperature rise can be too slow to

stress the package seals adequately.

The experimental apparatus consisted of a temperature-controlled hot stage

with the package placed on it face-down (as shown later in fig. 25 but with

the AE detector removed). The package temperature was monitored with a separate thermocouple. For this purpose, 0.003-in. diameter thermocouple wires

were welded to the inside wall of a hybrid package adjacent to and at the same level as the lead-throughs . In order to be sure the actual wall temperature was monitored, the thermocouple was split and each wire separately

welded, approximately 0.01 in. apart, to the package wall. Data were taken

at five-second intervals using a digital read-out thermometer. Data from the

first few runs showed poor reproducibility, so the hot plate surface was flattened and smoothed with a fine file. The package bottom was also lightly

polished. Figure 24, curve (a), gives the average temperature-time curve of

a package, open face up (the normal configuration), on the 400°C hot stage

for a gold-plated package of 1.25 by 1.25 by 0.3 in. dimensions. Each point

is the average of three runs. For comparison, curve (c) was taken under

conditions identical to (a) except that two short pieces of 0.003-in. diameter alumel thermocouple wire were placed on the hot stage under one edge of

the package. The wire was invisible except when viewed in oblique light.

Curve (d) was taken similar to (c) except that a 0.05-in. long by 0.015- by

0.010-in. piece of kovar from a package lead was placed underneath and near

one end of the package. The 0.01 5-in. dimension was parallel to the hot

stage surface. From examination of the curves, it is apparent that the initial temperature rise at the glass seal can vary considerably, and much care should be exercised in carrying out the test. Factors leading to varying

temperature rise at the seals are: package design (i.e., height of seals

above bottom surface and size of package), surface plating material and its

42
0 10 20 30 40 50 60 70

TIME |s|

Figure 24. Temperature-rise curves for a hybrid package on the hot stage of

figure 25. (a) Package face up, hot stage at 400°C (normal practice). (b)

Package face down, hot stage at 425°C. (c) Section of 0.003-in. wire under

face-up package (400°C) . (d) Section of package lead under face-up package

(400°C)

.
reflectivity, particles or other contaminants or irregularities on the

surface of the hot stage, and possible ambient air motion (drafts) across its

surface. "

After characterizing the normal hot stage thermal shock test, it was necessary to obtain equivalent data on the package face-down with convectionradiation shields and a cooled detector on the bottom. After several runs,

dt was determined that the equivalent rate of temperature rise in the

initial 20°C to 200°C range* was obtained with the stage temperature

'increased ~25°C to 425°C, as shown in figure 24, curve (b). However, because

this temperature exceeds that normally employed, a maximum temperature of

;400°C was set for further tests. This corresponds to a face-up shock

temperature of about 375°C which is still in the range normally used.

Only a single hot stage and two types of packages have been characterized in

these tests. Different organizations may use quite different hot stages,

ranging from small substrate heaters to large hot plates, as well as a large

variety of package types. Therefore, each particular setup should be

characterized, and the surface of the stage smoothed and kept clean in order

to obtain consistent results.

B-2 .

3

Acoustic-Emission-Monitored Shock Test

Most commercial hot plates generate AE during operation. For purposes of

acoustic-emission-monitored tests, a "quiet" hot plate is required. It was found that units heated by insert "slug" type elements meet this requirement,

whereas ones with wrapped coils, as are typical in commercial hot plates, are extremely noisy. One of the former designs was available from a heated stage

of a bonding machine and therefore all AE-monitored shock tests described

below used this type of hot stage.

The high temperatures of the hot stage required for this test presented dual

problems for AE detectors. Rapid heating of these detectors can produce

extraneous AE-like noise, and the most sensitive detectors have Curie point

restrictions and/or mount constructional materials (i.e., epoxy) that can be

damaged by the high temperatures. The first experiments to protect the

detector used a stand-off acoustical waveguide made of Pyrex attached to an

AE detector, the unattached end being acoustically coupled to the hot package

with an indium-gallium liquid alloy. The detection efficiency of this method

was very low because this waveguide extension had a low aperture for surface

waves, and in addition the indium-gallium attacked gold plating on the

package

.

Next, a small water-cooled detector mount was designed and constructed. This

adequately stabilized the detector temperature when the package was turned

face down and a small sheet of insulation material, such as silica

fiberglass, or woven ceramic, was interposed as a radiation-convection

barrier between the hot plate and the bottom of the package (lying within the

package cavity). A photograph of the cooled detector mount, hybrid package.

* This is the temperature range that produces maximum stress on the perimeter

of glass-metal seals in the hybrid configuration [40]

.

44
and hot stage is shown in figure 25. This particular mount used a commercial

AE detector. However, a similar mount was made for the TO-100 detectors.

Additional tests were made with a small thermocouple spot welded on the inside bottom of the package, directly under the detector. With the package

side walls in equilibrium (>1 min on a 400°C hot stage), the detector temperature cycled (in synchronization with the hot stage cycling) from 50°C to

54°C. With higher water flow through the cooler, the detector cycled between

43°C and 45°C. Either temperature range is satisfactory and does not lead to

spurious thermally induced signals. The water flow must be kept below the

turbulent range since turbulence generates large AE signals which can mask

the test results.

After temperature characteristics of the hot stage test and the cooling mount

of the AE detector were determined, an acoustic-emission-monitored study of

the thermal shock characteristics of various packages was carried out. A

block diagram of the general AE equipment used in section B of this report is

given in figure 26. Some tests may only use selected parts of the equipment.

AE output during the shock tests was anticipated to extend over a period of

up to 30 s, based upon the temperature-rise characteristics given in figure

24. Therefore, the apparatus capable of triggering on a single glass crack

was not necessary. Instead, just the detector in its water-cooled mount, the

totalizing system, and the printer portion of figure 26 was employed. Each

AE cycle or pulse above a set threshold was recorded and the summation of

these were printed each second. The plotted AE could then be compared with

the temperature rise versus time curves (fig. 24) to determine the average

temperature of the seal using the hot stage and AE detector apparatus of

figure 25.

Initially, a control package was fabricated from stainless steel. However,

because of residual machining strains as well as the high thermal expansion

coefficient (compared to kovar), this package was noisy and emitted frequent

AE bursts even after repeated tests on the 400°C hot stage. Similar tests on high-quality hybrid packages showed that they emitted essentially no AE during hot plate shock, so data from one of these were used as a control to

define a good package. The accumulated acoustic-emission counts* versus

time on the hot stage for the control sample are plotted as curve (a) in

figure 27. The small, essentially linear increase in signal with time is due

solely to the AE system noise above the set threshold and is not due to glass

seal degradation. This package and others were tested for leaks at room temperature and at 125°C. The control sample leak rate was <10~^ atm»cm^/s

at both temperatures.

Most of the 42 packages tested in this manner produced AE versus time curves similar to that of curve (a) of figure 27. However, a small lot of four

packages from one manufacturer that had been thought to be of high quality

showed large increases in AE when shock tested. Plots of the data from three

of these are given as curve (b) in figure 27 and curves (a) and (b) in figure

28. These figures show a dramatic increase in AE when compared to the con-

* AE counts as used here are the number of cycles of signal having amplitudes

above a set threshold. The threshold is usually set to minimize the

noise.

45
Figure 25. Thermal shock hot stage with a 1- by 1-in. hybrid package and a water-cooled AE detector on top. (a) Brass weight to hold the spring-loaded

detector against the package. (b) Part of the lead and connector from the AE

detector. (c) Water-cooling jacket around the AE detector. (d) Hybrid package shown face down on the hot stage. (e) Silicone rubber coolant tubing,

(f) Hot stage surrounded by a protective screen.

46
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49
trol sample data of curve (a) in figure 27. The delay of AE for a few seconds after placing the packages on the hot stage was expected, since from the

temperature rise curve (fig. 24) it takes several seconds to reach the 125°

to 150°C required to place the seals in tension. A retest of the high AEproducing samples resulted in very low AE output, similar to and only slightly greater than that of the control sample. This "one-time-only" phenomenon

is known as the Kaiser effect [2] in which little or no AE recurs until previously applied stress levels are exceeded. (Literally, a deformation or crack will only produce more AE if it is enlarged.

)

Shock tests were also run on several of the "bad" packages that had been vibration tested in section A. These packages had undergone liquid-to-liquid

shock, and many of the glass seals had previously become leakers. Very few

of the leaking packages produced large AE signals. However, one of the unvibrated samples from the bad lot that was tested and found to be hermetic did

produce large AE in the hot stage shock test similar to that shown in figure

28. The five packages that produced large AE signals were leak tested both

at room temperature and at ~120°C. All of them were hermetic (<10~^

atm»cm^/s) at room temperature. However, two of these packages became leakers (>10~^ atm»cm^/s) at ~120°C, and the seals that leaked were identified by

carefully spraying helium over the glass beads. There was no visual damage

to identify the leaking seals. This temperature, 120°C, is about the minimum

at which the forces in the perimeter of kovar glass seals change from com- pression to tension. The leak testing apparatus in its current configuration

(see sec. B-1 ) cannot go to a higher temperature. However, since the AE test

detected measurable seal damage in all five of these packages, it is speculated that the other three packages would also leak if the test temperature

could have been raised somewhat higher. Even if the other three remained

hermetic, a 40-percent detection rate for marginal packages is adequate to

reveal seal problems on a lot basis. The total number of packages tested in

this experiment was relatively small. Only 48 packages from three manufacturers involving six production lots were available for this particular

study. It is recommended that a larger study, perhaps involving direct manufacturer participation, be undertaken to verify the results of the present

work and to develop practical procedures for test implementation.

The AE shock test configuration used for flatpacks cannot be used for testing

plug-in-s tyle packages, and no such tests were performed on them in this

work. The seals (and leads) in plug-in packages are on the bottom, significantly removed from the side walls that heat rapidly when the package is

placed face down on a hot stage. A suggested configuration for AE-monitored

shock testing of plug-in packages would incorporate a metal jig on the hot

stage that would be similar in shape to a package socket but with larger lead

pin holes to allow clearance of the glass beads. The AE detector could be

attached to a normal package lid. The lid could be acoustically coupled to

the package with a high-temperature vacuum grease, which would be cleaned off

later. Because of the fast heat transfer to the glass leads in this configuration, the hot stage temperature might be lowered to perhaps 250°C and

still match the 0°C to 150°C temperature rise given in the curves of figure

24.

If large-scale verification of the present results is obtained, AE shock

tests can be effectively performed at the package manufacturer's plant on a

50
lot sample basis to assure delivery of only high-quality lots. The verified Kaiser effect for this test would render results obtained on incoming

inspection at a user's facility somewhat less definitive unless the thermal

shock history of the packages were known. Without such information, a hightemperature open-package leak test might be adequate, although the AE test

could be used for added assurance.

The AE-monitored shock test can be performed more rapidly (about 10 to 15 s)

and economically than a liquid-liquid shock test followed by a leak test.

The ability of this test method to reveal marginal packages is unique and

therefore warrants verification on a large scale. It would also be interesting to correlate AE tests with high temperature die penetrant tests. AE

monitoring might also be used with shock tests at lower temperatures (~150°C)

for hybrid component epoxy or solder bond integrity as well as to reveal

substrate cracks.

B-2 .

4

The Possible Use of a Thermoelectric Temperature Cycler for Stressing

Hybrid Packages in Conjunction with Acoustic-Emission Monitoring

A thermoelectric temperature-cycling system has been used in these studies to

stress glass-to-metal seals in hybrid packages repeatedly from -55°C to

125°C. It had been expected that such a system would be acoustically "quiet"

and that the packages could be monitored for acoustic emission during rapid

cycling to supplement and extend the hot stage shock test. However, it became apparent that the thermoelectric modules themselves emitted large bursts

of AE, especially when they were cooling. There were several possible ex- planations. One was that the polycrys talline bismuth-tellurium alloys, which

make up the thermoelectric cooling elements, underwent crystallographic

changes such as twinning. Others concerned possible twinning in the solder

as well as crack propagation in the numerous interfaces between the solder

and the Bi-Te segments or in the Bi-Te material itself. The latter possibilities might be curable by using different assembly procedures and materials.

However, if the problem resulted from the Bi-Te material twinning, then thermoelectric temperature cyclers using Bi-Te elements could not be used for

stressing packages that were being monitored with acoustic-emission sensors.

Slices from polycrys talline Bi-Te boules, of both n- and p-type, were obtained for test. Three experiments were performed: (1) slow increase in

temperature as a hot stage heated up, (2) rapid surface heating by a concentrated infrared source, and (3) pressure. An acoustic-emission detector was coupled to the test samples at all times. Individual AE events were captured

with a transient recorder and counts per second were recorded using the apparatus in figure 26.

The Bi-Te boule was placed on a "quiet" heat stage (shown in fig. 25) and

further isolated from the stage with a ~1/16-in. thickness of silicone rubber. For the test, the temperature of the boule was increased at a rate of

~5°C per minute. Large acoustic-emission bursts were observed on both heating and cooling. The average number of AE counts recorded was in the range

of 200 to 300 per second (the background was 3 to 5 counts per second).

Next, the side of the boule was heated with a 500 W focused projector lamp

placed 6 in. from the sample. This created a temperature-differentialinduced strain near the surface similar to that of a thermoelectric cooling

51
element under power. Numerous large acoustic-emission bursts were observed

within 1 s of turning the lamp on. AE counts of 3000 to 4000 per second were recorded for about 20 to 30 s until the average sample temperature had risen

by about 30°C, then the count rate decreased to about 1000 per second. An

example of an AE burst occurring during the rapid heating period is given in

figure 29a.

Twinning and other crystallographic changes can often be induced by pressure

in susceptible materials. Therefore, a simple pressure experiment was run. Force was applied by a "quiet" piece of tapered nylon with a contact surface

area of 0.006 cm^; AE bursts were usually observed to initiate in the 400- to

1 000-gf range (64 to 160 kg/cm^), although there were considerable variations

over the surface of the sample and the number of bursts in a given pressure

range was not uniform. Figure 29b is an AE burst obtained at 64 kg/cm^.

Bursts were also obtained during force unloading, sometimes at forces as low

as 10 kg/cm^.

In addition to acoustic-emission studies, several thermoelectric modules that

had been cycled more than 1000 times and had become inefficient on the cooling cycle were visually examined around their Bi-Te element solder joints.

Microcracks were observed in the solder. Also surface reconstruction of the

solder was evident, indicating recrys tallization had taken place, and finally

some cracks in the Bi-Te elements themselves were observed.

All four possible sources of AE in thermoelectric modules have been verified

to exist; therefore, it does not appear possible to use thermoelectric

temperature-cycling equipment in electronics or other applications that re- quire acoustic-emission monitoring.*

B-3 Determination of the Susceptibility of Kovar Glass Seals

to Damage During Wire Bonding

B-3 .

1

Mechanical Characteristics of Kovar Leads in Hybrid Packages

There are a number of stresses that are applied to hybrid package seals during production that may damage those seals. Two obvious examples are lead

forming (bending) and wire bonding to the internal package extension of the

leads. This section describes work directed toward determining the levels of

stress encountered during wire bonding that may damage the seals and, if

found to be destructive, recommending procedural or other changes to retain

seal integrity. The approach used was to determine mechanical characteristics by "beam bending" the lead extension inside the package, and any glassseal cracking would be revealed by an AE sensor coupled to the package.

* It should be noted that most of the thermoelectric module degradation

processes, such as cracks, are considered to result from the large range of

temperature cycling (-55°C to 125°C). Thermoelectric modules are generally

used only for cooling below room temperature; thus crack propagation and

solder reconstruction would probably not be significant under normal usage.

Twinning of the Bi-Te, which would occur during normal cooling, should not

degrade the thermoelectric performance.

52
(a)

(b)

Figure 29. Acoustic-emission bursts resulting from stressing Bi-Te boules.

(a) Burst obtained during rapid heating of a boule. The burst is attributed

to twinning. (b) Burst resulting from the application of a force of 54 kg/cm'

to the surface of a boule. The horizontal scale in (a) and (b) is 200 ys per

major division.

53
Figure 30. Sketch identifying beam-bending variables.

The simplified cantilever beam equation, in which the entire bending force is

applied to one end while the other end is held fixed, is given in eq (3).

The variables are identified in figure 30.

where:

y

= -F£3 -4

3 EI bd^E (3)

F = applied force, lbs.;

Z = the beam length, in.;

I - 1/12 bd^ ~ section moment of inertia; and

E = Young's modulus -19.8 x 10^ psi for kovar [40].

The general equation of deflection at any point,

with the force applied at any point, L [44], is:

X, along a cantilever beam

y

= _FXj_

2b ^ E

(X - 3L) - FL^

3bd^E ix - L)Q

when X > L, Q = 1 when X < L, Q = 0

(4)

also X = X when x < L, and X = L when x > L; and

L = point of force application (in. from beam attachment).

It should be stressed that deflection is proportional to the cube of the lead

length and inversely proportional to the cube of the lead thickness. Thus,

it is important that the package manufacturer control these dimensions to

minimize lead bending and any effect it can have on wire bonding, as well as for compliance with lead integrity tests (MIL-STD-883 Method 2004.2) [36].

The ISHM Hybrid Microelectronic Specification for Metal Packages (ISHM MSP008, 1978) [45] and most manufacturers' catalogs allow a ±0.002-in. variation in lead thickness. This allowed variation can significantly affect

bending under a given force as indicated in eqs (3) and (4). Therefore, lead

thickness measurements were made on a number of the packages at hand in order

to determine the typical range of lead thicknesses actually delivered. These

75 packages were predominantly from two manufacturers and included samples

from many lots. The thicknesses measured are given in table 3. 54
Table 3. Measured Thicknesses of Hybrid Package Leads.

Specified

Thickness, in. 0.005 in. 0.01 in.

Measured

Thickness, in.

0.0045 0.005 0.008 0.009 0.0095 0.010 0.0105 0.01 1

Number of

Packages

6 2 1 3 1 4 38 10 1

In a few cases leads of the same package ranged in thickness from 0.009 to

0.011 in. (within tolerance) and in width from 0.011 to 0.016 in. (uncontrolled). Although the lead thicknesses were all within tolerance, there is an obvious skewing of the distribution towards the thin side where bending is

easier. Packages from more manufacturers might show a larger spread.

The depth of lead extension into the package, where wire bonding takes place,

is equally critical with lead thickness for bending under a given force. The

packages studied showed interior lead lengths of from 0,025 to 0.054 in.

Packages can be obtained from manufacturers with various internal lead extensions ranging from 0.025 in. to 0.110 in. ^q' qq^

• Long leads will bend significantly even under the low forces of ultrasonic wedge bonding, presenting

nonflat bonding surfaces which could result in relatively unreliable bonds.

The bending of long leads would be severe under thermocompression (TC) or thermosonic (TS) bonding forces. If long leads (>0.05 in.) are encountered

in a package, wire bonding should take place close to the glass bead rather

than near the end of the lead as is the normal practice.

B-3 .Characteristics of Kovar Leads under Simulated Wire Bonding Stresses

Tests were run to determine the characteristics of leads under thermocompression and thermosonic bonding forces. The apparatus used in these leadbending experiments is shown in figure 31, with a closeup of the force probe,

package, and AE detector shown in figure 32. Measurement of the deflection

was made with a 1 20X binocular microscope having an eyepiece reticle which

was calibrated with a precision stage micrometer. The force was applied by a special thin bonding tool and measured by the force gage.

Initial lead-bending tests were run on packages with relatively long leads in

order to determine the mechanical characteristics of the kovar more easily.

The leads of the packages selected were 0.010 in. thick by 0.014 in. wide and

0.10 in. long. The actual deflection was measured and then calculated using

the published kovar modulus in the equation. The results of one such test

are given in table 4.

55
CL) 1 o

Xi T3 o Q C0 1 o ,

X3

TO

Q) (Ti r;1 1 o

0) 0)

M-l X! 0 in O

Cn

C

•H Cl, 0

•H <DO C 0 •H 0 C X o

(1) ti he c 0

a> of

0) 0

-P ro p u 0 (U

CU 0 4-1 !h U) o -p 0) 3

Cn 13

u Cn

H

C

C

03 < •H

O rO

4-1 0)

OJ

H

4-1 M

O

03

03

U

0)

O

U 4-1

4-1

C

•H < ^3CO • Xi

CO 01

4-1 01 03

!^ -H

56
Table 4. Comparison Between Measured and Calculated Bending of Kovar Leads

for Leads of 0.01 x 0.014 x 0.10 in.

Force (g) Measured Deflect, Calculated Deflect, Ratio

in. in. (meas /calc

)

9 0.66 X 10-3 0.28 X 10-3 2.36

18 1 .33 0.57 2.33

25.5 1 .99 0.80 2.52

34 2.65 1 .07 2.48

36 3.3* 1.13 2.92

* yielding observed.

The large differences (ratios) between experimental and calculated bending

deflections have been verified on a number of different packages and with

leads of different lengths and thicknesses, both with and without gold plating. The discrepancy cannot be attributed to yielding since the beam deflection returned to zero upon unloading for the first four force values in table

4, and the changes in the ratio (2.36 to 2.52) are within measurement error. However, the highest force was observed to bend the lead permanently. The

only material constant in eq (4) is the modulus which is 19.8 x 10^ psi for

kovar. One would not expect any possible small compositional changes in the

kovar to change the modulus enough to account for the discrepancy.

There may be a number of possible explanations for the discrepancy. However,

the forces on the glass under the beam may be as high as 40,000 psi, and

compression of the glass is possible.

An analysis of the contribution to lead bending resulting from the elasticity

of the glass bead is based on the following equation for the slope of a beam

attached to a compressible (nonrigid) support [46].

16.67 M ( 1 - a) V

Ed'^ Ed

where the variables, not elsewhere defined, are:

K, the slope of the beam,

M, the bending moment (in. -lb. /in. )

,

V, the shear force per unit width of beam at the support (lbs. /in.),

E, the modulus for a typical kovar sealing glass, 8.2 x 10^ psi,*

a, Poisson's ratio for kovar sealing glass, 0.24.

The actual beam deflection due to compression of the elastic support is

The deflection calculated from eq (5) and multiplied by i is added to the

deflection calculated from eq (4) to obtain the total deflection. Table 5 gives the corrected calculations and the ratio between calculated and measured deflection.

* Such as for Corning 7052 glass.

57
Table 5. Kovar Lead Bending Corrected for Glass Elasticity.

Force

(g)

Measured Deflection

(in. ) Calculated Deflect

Eq (4) + (£K)

(in. )

Ratio

(meas . /calc.

)

9 0.66 X 10~3 0.57 x 10"^ 1.15

18 1.33 1.14 1.16

25.5 1.99 1.62 1.23

34 2.65 2.16 1.23

36 3.3* 2.29 1.44*

* yielding observed.

Additional corrections for the glass meniscus and bending of the package

kovar wall could lead to an even closer correlation between the measured and

the calculated values. However, the results of table 5 bring the difference

between the measured and calculated deflection within reasonable agreement

and illustrate the importance of including all portions of the glass -metal

seal when making a stress analysis of hybrid packages. In addition, these

tests show that the often-made assumption that glass can be considered rigid

in small seals is incorrect.

B-3 .

3

Acoustic-Emission-Monitored Tests to Determine Glass-Seal Damage

During Bonding

B-3. 3.1 Room-Temperature -Bonding Experiments

After determining the mechanical characteristics of kovar leads in sections

B-3.1 and B-3. 2, acoustic-emission-instrumented tests were begun in order to

establish whether glass-seal cracking could be induced by the forces and

temperatures of thermocompression and thermosonic bonding to the kovar leads

and to establish whether any such cracking could be detected with the TO-100

acoustic-emission detector, which was waxed to the hybrid package, as was shown in the closeup of figure 32. (The block diagram of the AE circuit was given in fig. 26. ) Both the transient recorder and the counts-per-second

totalizer were used in this experiment.

Initial room temperature tests began on packages having relatively short lead

extensions (~0.035 in.) inside the packages. The force was continuously

monitored and recorded each time an AE burst occurred. The first observation

was that the leads of certain packages produced numerous AE bursts in the

relatively low force range, £50 gf. One such AE burst at a 45-gf load,

shown in figure 33, was traced to a crack in the glass under the lead. Another larger burst that saturated the amplifier is shown in figure 34. In

this case a sliver of glass under the lead cracked off at only a 35-gf load.

One lead emitted bursts at 19, 39, and 60 gf. Packages that emitted bursts

under low loads were observed to have glass wetting or "wicking" extending

irregularly outward on the lead past the normal meniscus of the glass bead.

This thin glass can crack off with minimal lead bending. There is no evidence thus far that such low-force-induced cracks affect the hermeticity

58
Figure 33. Acoustic-emission burst from glass cracking due to a force of 45

gf applied to the internal lead. The horizontal scale is 220 ys/div.

Figure 34. A large acoustic-emission burst (saturating the transient recorder)

resulting from a small glass sliver cracking off from the bottom side of the

lead and becoming a loose particle. The horizontal scale is 220 ys/div.

59
of the seal. However, glass slivers released (or cracked, but left loosely

attached during the bending) could certainly act as loose particles if not

removed before package sealing. The cracking and slivers generally occur in

the package on the underside of the lead where it is difficult or impossible

to observe. As a result, it is possible that such cracked or broken glass

could be a significant source of nonmetallic loose particles in susceptible

packages . Package specifications should require that the inside glass bead

be either flat or have a small uniform meniscus shape.

An unusual form of protection against uneven miniscus glass cracking was observed. One step in the manufacture of packages, after fusing the glass

beads and sealing the leads in the package, involves etch removal of the excess kovar oxide. If a package remains in the etch bath too long or is not

completely cleaned, undercutting of the kovar below the glass may occur. Such undercutting is not normally considered desirable and if severe could

lower the bending yield point of the lead. However, the combination of the

slightly thinned kovar and the resulting clearance between the most extended

part of the glass meniscus apparently serves as protection against glass

cracking during slight bending of the lead. Therefore, even though the visual appearance of etch undercut leads is not ideal, such etching appears to

protect the glass against cracking under relatively light-load bending. It

should be pointed out, however, that when such etch undercut patterns appear

on the outside of the packages, they could serve as stress risers or weak

points and result in early fatigue damage in vibration environments (sec.

A).

Several additional experiments were carried out using higher bending forces

on the leads to determine the level required to produce significant glassseal damage. The results indicate that, on the average, significant glass

miniscus cracking occurs in packages with 0.010-in. thick leads extending

inside the packages approximately 0.035 in., at forces between 150 gf and

200 gf. However, severe glass cracking was detected both by AE and visually

on one seal with a very irregular meniscus at a force of 125 gf.

Other tests were carried out on seals that had been significantly cracked on

the outside by intentionally bending the leads. These seals were subjected

to probe forces on the lead extension inside the package as described above

using the apparatus of figure 31. One such lead produced large AE bursts at

40, 95, 125, and 200 gf. These bursts were in the 200 to 500 count range and

of sufficient amplitude to saturate the transient recorder input with no evidence of an inside sliver cracking off. Another seal that was less damaged than the above one produced several unusual double bursts when stressed,

an example of which is shown in figure 35. Presumably, a small crack propagated into a larger one. The total AE count in this burst was 367 as registered by the totalizer. Apparently, in both of the above cases, the force

was transmitted through the seal and resulted in propagating already formed

cracks on the outside. It is reasonable to expect that force can be transmitted through a glass seal since the glass is compressible and has a relatively low modulus of 8.2 x 10^ psi. This compressibility was discussed in

section B-3.2.

60
Figure 35, An AE burst at 25 gf applied to the lead from a seal that was damaged on the outside with no damage observed on the inner side of the seal.

The double burst presumably resulted from a small glass crack propagating and

causing a larger crack. Seals that are damaged on the outside appear to be

susceptible to additional glass cracking under the stress of bonding forces.

The horizontal scale is 220 ys/div.

61
B-3 .

4

AE-Monitored Package Lead Bending under Thermosonic Bonding

Conditions

After the room-temperature bending characteristics of kovar leads were established on special laboratory equipment (fig. 31), the experiment was moved to

a bonding machine with a heated stage, where any damage to the seals could be

determined under bonding conditions. A photograph of the apparatus, including part of the bonding machine, is shown in figure 36. For convenience, the

force was applied on the external lead extension at a fixed distance from the

glass bead (usually 0.035 in.). The intent was to reproduce the experiments

of section B-3. 3 but at thermosonic bonding temperatures (~125°C to 150°C)

and with the possibility of adding ultrasonic energy during the application

of force. Interference between the 60-kHz ultrasonic bonding signal and

possible acoustic emission was anticipated so the AE-detector output was passed through a 24-dB per octave tunable filter plus appropriate 1 8-dB per

octave fixed frequency filters to minimize such interference. Actual wire

bonding to the leads was performed to determine the frequency and bandwidth

of the required filtering. To sharpen the response peak further, measurements were made at the sensitivity peaks of the TO-100 AE detector (see fig.

15). Because of the very large amplitude of the 60-kHz fundamental produced

by the bonder as well as the numerous harmonics generated during bond formation, there was always too much interference to separate small glass cracking

from the extraneous "noise." Only in the case of very large induced glass

cracks could the AE signal be distinguished at all, and clear separation was not possible.

The result of this work revealed that there is an empirical relationship

between various stages of bond formation and the AE-detector output and thus

this technique might be useful as a bond monitor. Three waveforms are shown

to illustrate the point. Figure 37 shows approximately the first 10 percent

of a 50-ms bonding cycle. The individual "spikes" occur at the 60-kHz fundamental of the bonding machine. Figure 38 is a 5X expansion of the waveform

starting at the arrow in figure 37. Figure 39 is the beginning of a second

ultrasonic bonding pulse on the same (now mature) bond started in figure 37,

The 60-kHz "spikes" are gone and the amplitude of the envelope varies along

the bonding pulse. These three photographs were taken with the two AE-output

filters placed in series and set for a bandpass of 650 to 800 kHz, and this

response was further narrowed by the AE detector response. From this and the

detector frequency response of figure 15, it is apparent that most of the

signal was either the 11th or 12th harmonic of the 60-kHz ultrasonic bonding

pulse. In any future studies of this phenomenon, it would be desirable to use a flat-response broad-band AE detector and perform an FFT on the signal

in order to determine its usefulness as a bond monitor. One should also

consider the possibility that the signal amplitude and frequency characteristics could be dependent upon the package geometry and the position of the AE

detector.

Since it was impractical to monitor AE generated by hermetic seal cracks that

occurred during ultrasonic or thermosonic bonding, the investigation was limited to studying glass cracking that resulted from the equivalent combination of temperature and force but without ultrasonic energy which might

simulate thermocompression bonding. The equipment was as shown in figure 36,

except that a thin, 60-deg bonding tool was used as the probe in order to

62
Figure 36. Apparatus to apply thermosonic- type forces and monitor any seal

cracking in hybrid packages. (1) Acoustic-emission detector in a water-cooled

mount. (2) Hybrid package face down with its external leads cut short. (3)

Substrate heater to produce thermosonic bonding temperatures. (4) Bonding

machine ultrasonic transducer and bonding tool.

63
Figure 37. The first 10 percent of a 50-ms ultrasonic bonding pulse as detected by the AE detector in the equipment setup of figure 35. The "spikes"

appear at 60-kHz time intervals. The substrate temperature was 12 5°C. This

entire ultrasonic pulse resulted in bonding a 2-mil diameter aluminiam wire to

a hybrid package lead at approximately 0.03 in. from the glass seal. The

horizontal scale is 36 us per major division.

Figure 38. A 5X expansion of the waveform in figure 37 centered on the position of the arrow.

64
Figure 39. The initial 8 percent of a second 60-kHz bonding pulse applied to

the same bond that was completely formed by the pulse shown in figure 37. This

second pulse was applied immediately after the first, and the bonding tool re- mained in contact with the top of the bond. The 60-kHz "spikes" are gone, but

the periodic bunching of the waveform is still in evidence. The horizontal

scale is somewhat expanded from that of figure 36 (27 ys per major division)

65
apply force nearer the glass seal. The bondinq machine was mechanically

cycled in order to include the results of any impact loading. Substrate

temperatures in various tests ranged from 100°C to 175°C and forces up to 200

gf. It was considered possible that the change in the glass-metal-seal internal forces from compression to tension at these temperatures might result

in more damage than was observed during similar room temperature tests. The

results of applying such stresses to 10 leads from each of 14 packages indicated that there is no significant difference between the force required to

produce glass cracking at high or at room temperature. The high-temperature

stressing appeared to result in a somewhat greater number of AE bursts (glass

cracking), but the variations from package to package made the increases

statistically insignificant. Most of the bursts were small, from 50 to 200

counts, and presumably resulted from tiny meniscus cracks that are of no significance as far as package integrity is concerned. The package hermeticity during these tests was affected in only one qualified case.* Even

the breaking of glass slivers which produced large bursts of several hundred

AE counts did not affect hermeticity. However, these slivers may result in

loose particles, which could cause rejection or rework of expensive finished

hybrids

.

Most hybrid packages have relatively short lead extensions into the package.

For 0.010- by 0.01 5-in. rectangular leads, this extension typically ranges

from 0.030 to 0.048 in. No significant bending during bonding with normal

forces, up to 125 gf, has been observed on such leads. However, forces on

the order of 150 gf applied near the tip of all but the shortest leads will

permanently bend them downward a significant amount and thus may result in

weak bond formation. Such forces may be encountered if the bonding machine

has a significant impact loading or when bonding with large wire (>0.001-in.

diameter). The longer or thinner the lead, the lower the force required to

bend it and, therefore, the greater the chance to produce a poorly welded

bond.

B-3.

5

Metallurgical Properties of Hybrid Package Leads^

Six flatpack hybrid packages with leads extending out of two sides were evaluated for the hardness and plating thickness of their leads. All came from

the same manufacturer but were from three production lots made over a period

of about a year. An evaluation was made of the hardness of the leads on each

of the samples to determine if substantial variation existed in the lead

strength of the three lots of kovar. The leads were approximately 0.01 in.

thick and 0.015 in. wide with various thicknesses of gold and nickel plating.

* One package from the GIDEP Alert series [38] (see A-3.1) that initially had

a fine leak did change to a gross leak after a number of its leads were bent at 200 gf. The AE bursts were similar to those from other packages

subject to similar forces. It is not known whether this failure resulted

from the mechanical unplugging of a gross leak or the creation of a new

one. The package had previously been used in another test and may in some way have been partially damaged. No othf - "good" packages from reject lots were available to confirm this failure.

t These measurements were supplied by Dr. John H. Smith, NBS Fracture and

Deformation Division.

66
Leads were cut from each of the six samples and mounted in metallographic

mounting material. Vickers hardness readings [ASTM E 92-72 (1977)] [42] were made on the ends of the leads using a 500-g load. Hardness measurements

were made on three leads from each end of the samples; that is, six leads were measured from each of the six samples. The results showed that all 36

hardness values were within the range of 144 to 150 on the Vickers scale (HV

0.5) using a 500-g load. This variation is not significant, so it is con- cluded that all six samples have the same hardness, and therefore, it is

inferred that they all have the same strength.

Photomicrographs were taken of leads from each of the six samples evaluated.

These photographs are shown in figure 40. Measurements of the thickness of

the plating on each sample were made from these photomicrographs and are given in the figure captions.

Additional hardness measurements were made on the sides of the leads from

samples shown in figures 40d and 40f. This was done to evaluate the possibility of using hardness measurements for inspection of incoming material.

Hardness measurements were made using 500-, 1000-, and 2000-g loads on the

Vickers hardness machine. The hardness results on the sample in figure 40d,

which was unplated kovar, ranged from 148 to 158 HV. The hardness measurements taken on the sample in figure 40f, which was plated with both nickel

and gold, showed wide variations due to the presence of the plating. The

measured hardness varied from 130 to 176 HV with a 500-g load and from 370 to

850 HV with the 2000-g load. These results show that the plating interferes

with the measurement and makes it impossible to determine the kovar hardness

on plated specimens. The plating should be polished off before such a test

could be useful. On the basis of this limited test of lead hardness on samples from one manufacturer, hardness testing does not appear to be a useful

screen for kovar lead quality. No measurements were made on the package

casing, which could be a more variable material.

B-4 Conclusions of Part B

Part B of this report concerned the development of several types of nondestructive tests to assess the integrity of hybrid packages.

A 125°C open package leak test method was developed along with the proper

procedures for handling and preconditioning any type of hybrid package for

hermeticity testing.

The hot plate thermal shock test often used by package manufacturers to assess the quality of glass-metal seals was characterized and various problems

and pitfalls revealed. This quick, simple test, when monitored by acousticemission equipment, can reveal marginal seals that could only be detected by

high-temperature leak tests. A thermoelectric temperature-cooling apparatus

was evaluated for possible use as a "quiet" temperature-cycling system com- patible with acoustic-emission monitoring of thermal-cycle-induced failures

in hybrids. It was found that the thermoelectric modules themselves produced

large amounts of acoustic emission which is fundamental to the materials and

construction methods. No means of acoustic isolation that would still allow

rapid heat transfer was found. Therefore, this means of temperature cycling

cannot be used for acoustic-emission-monitored experiments.

67
(a) 0.0004-in. nickel under 0.0001-in. (d) No coating,

gold coating.

(b) 0.00025-in. gold coating. (e) 0.0001-in. gold coating.

(c) 0.0002-in. gold coating. (f) 0.00015-in. nickel under 0.0003-in.

gold coating.

Figure 40. Photomicrographs of kovar leads from six different hybrid packages. Each

package had a different surface coating (plating). (Magnification, 750X - as polished)

68
A study was carried out to determine the susceptibility of the glass-metal

seals in hybrids to damage during wire bonding. It was found that the amount

of lead bending under bonding loads is significantly greater than predicted

by theory, due to an extra component contributed by the compressibility of

the glass. It was also found that the thickness of kovar leads tended to be

skewed towards thinner-than-specified dimensions. This could result in increased bending and less reliable bonding. Under some bonding-force and

glass-wetting conditions, glass slivers can be loosened or cracked off the

kovar lead and could become loose particles after the package is sealed.

Acoustic-emission-monitored tests to assess possible seal damage during thermosonic bonding could not be successfully carried out because any acoustic

emission due to glass cracking was masked by harmonics of the ultrasonic

bonding pulse. The AE detector did sense signals that appear to be related

to bond formation, and thus this technique may be useful as an in-process

bond monitor. Measurements of glass-seal damage monitored by AE were made at

thermosonic bonding temperatures of up to 175°C and forces to 200 g. Results

revealed that the seals on high-quality packages are not significantly more susceptible to damage at these high temperatures than at room temperature.

Forces on the order of 150 gf, applied to the ends of typical internal lead

extensions, will cause significant bending which can result in poor bond

formation.

The major problem in this study was to obtain packages from failure-prone

lots that were still initially hermetic. Even with the generous cooperation

of two package manufacturers and several user organizations, there were never enough "poor" packages to prove definitively the utility of all of the test

methods. Future work should be preceded by an industry-wide appeal for appropriate specimens so that a statistically significant sample can be accumulated. The general conclusion of this work is that small meniscus glass

cracks have no effect on the hermeticity of the package and that it is difficult to produce hermetic damage to the seals of high-quality packages. However, packages screened and found to be "good" from poor or reject lots (lots

with a high percentage of failures) were found to be more susceptible to

hermetic failure from all types of stress tests than packages from highquality lots. An acoustic-emission-monitored shock test has been shown to

separate "good" from "poor" package lots quickly. Additional verification of

these results with a statistically significant number of packages is recommended.

69
Acknowledgments

K. A. Harmison, Naval Avionics Center, supplied some hybrid packages and

printed circuit boards as well as information and cooperation essential to

this work. The author would like to acknowledge valuable discussions with H. E. Hagy of Corning Glass. In addition, some special measurements in this

report were performed by Drs. N. N. Hsu and F. R. Breckenridge, Mechanical

Production Metrology Division, NBS . Stanley Ruthberg made all leak test

measurements and contributed sections B-1 and B-2. Section B-3.5 was con- tributed by Dr. John H. Smith of the Fracture and Deformation Division, NBS.

Cooperation from several hybrid package manufacturers, notably Tekform and

Isotronics, is gratefully acknowledged. The manuscript was prepared by ^4rs.

Jane Walters.

70
REFERENCES

1. Harman, G. G., The Use of Acoustic Emission to Determine the Integrity

of Large Hybrid Packages, NBSIR 80-2055 (June 1980).

2. Kaiser, J., Untersuchungen uber das Auftreten von Geraschen Beim Zugversuch, Avkiv fur das EisenhuttemJesen 24_ (1/2), 43-45 (1953).

3. Spanner, J. C, Aaoustic Emission Techniques and Applications (Society for

Nondestructive Testing, Columbus, Ohio, 1974).

4. For a compilation of articles on AE, see Acoustic Emission, ASTM Technical

Publication STP 505 (American Society of Testing and Materials, Philadelphia, Pennsylvania, 1972).

5. The Western Electric Engineer 23_, 3-37 ( 1979). includes four papery on AE testing.

6. Vahaviolos, S. J., Real Time Detection of Microcracks in Brittle Materials Using Stress Wave Emission, IEEE Trans. Parts, Hybrids, and Packaging ^HP^^, 152-159 (1974).

7. Saifi, M. A., and Vahaviolos, S. J., Laser Spot Welding and Real-Time

Evaluation, IEEE J. Quantum Electronics QE-1

, 129-136 (1976).

8. Carlos, M. F., and Jon, M. C. , Detection and Cracking During Rotational

Soldering of a High Reliability and Voltage Ceramic Capacitor, Proc

28th IEEE Electronic Components Conference, Anaheim, California, April

24-26, 1978, pp. 336-339.

9. Jon, M. C. , Duncan, H. A., and Vahaviolos, S. J., Analysis of Stress

Wave Emission in Resistance Welding of Tantalum Capacitor, Materials

Evaluation 36_, 40-44 (1978).

10. Jon, M. C. , Keskimaki, C. A., and Vahaviolos, S. J., Testing Resistance

Spot Welds Using Stress Wave Emission Techniques, Materials Evaluation

36 , 41-51 (1978).

11. Knollman, G. C, and Weaver, J. L. , Evaluation of an Acoustic Emission

Monitor for On-Line Quality Control in Spotwelding of Electronic Components, Proc. Third Acoustic Emission Symposium, Tokyo, Japan, September 16-18, 1976, pp. 413-427.

12. Ikoma, T. , Ogura, M. , and Adachi, Y. , Acoustic Emission from Single

Crystals of Gallium Arsenide, Ibid., pp. 329-341.

13. Harman, G. G. , The Use of Acoustic Emission in a Test for Beam Lead,

TAB, and Hybrid Chip Capacitor Bond Integrity, Proc. 14th Annual IEEE

Reliability Physics Symposium, Las Vegas, Nevada, April 20-22, 1976,

pp. 86-98.

14. Kumar, A., Failure Analysis of Electric Component Leads, Proc. ATFA1979, Los Angeles, California, October 8-11, 1979, pp. 28-33.

71
15. Steinberg, D, S., Vibration Analysis fov Eleetvonia Equipment (John

Wiley & Sons, New York, 1973).

16. Steinberg, D. S., Circuit Components vs. Random Vibration, Eleetvonia

Packaging and Production, 185-187 (July 1975).

17. Steinberg, D. S., Avoiding Vibration in Odd-Shaped Printed Circuit

Boards, Machine Design 48_, 116-119 (May 1976).

18. Steinberg, D. S., Preventing Vibration Damage in Electronic Assemblies,

Machine Design 48_, 74-77 (July 1976).

19. Steinberg, D. S., Avoid Failure of PC Board Components, Electronic

Design 22_, 68-70 (October 1976).

20. Steinberg, D. S., Snubbers Calm PC Board Vibration, Machine Design

49, 71-73 (March 1977).

21 . Steinberg, D. S, Countering the Effects of Vibration on Chassis Systems,

Electronics 50, 100-102 (August 1977).

22. Sevy, R. W. , and Earls, D. L. , The Prediction of Internal Vibration

Levels of Flight Vehicle Equipment (Limited), Shock and Vibration

Bulletin 38 (Supplement), 5-18 (1968).

23. Kohler, B. A., Resonant Beam High "G" Vibration Testing, Shock and

Vibration Bulletin 39 (2), 171-176 (1969).

24. Isada, N. M. , and Shear, J. C, Vibratory Response of Printed Circuit

Boards, Shock and Vibration Bulletin 40 (3), 111-118 (1969).

25. Palmisano, R. R., and Neily, D. W. , Particulate Silicone Rubber: An

Effective, Removable Encapsulant for Electronic Packaging, Shock and

Vibration Bulletin 46 (4), 277-283 (1976).

26. Blanks, H. S., Accelerated Fatigue Life Testing of Leads and Soldered

Joints, Microelectronics and Reliability J_5, 21 3-219 ( 1 976).

27. Morse, P. M., Vibration and Sound (McGraw-Hill Book Co., Inc., New York,

1948).

28. Warburton, G. D. , The Vibration of Rectangular Plates, Proc, Instr,

Mech. Engr. 1 68 , 371-380 (1954).

29. Hearmon, R. F. S., Communications on the Vibration of Rectangular

Plates, Proc. Instr. Mech. Engr. 1 68 , 381-384 (1954).

30. Hearmon, R. F. S., The Frequency of Vibration of Rectangular Isotopic

Plates, J. Appl. Mechanics ASME T4_, 402-415 (1 952).

31. Rayleigh, Lord, Theory of Sound, Vol. 1 (MacMillan s Co., London, 1894).

72
32. Nelson, F. C, Damping PC Boards with Foam Plastic, Circuits Manufa&tuving ^9_, 50-53 (November 1979). (This paper gives several other ref- erences not included in the present study.

33. Curtis, A. J,, Tinling, N. G. , and Abstein, H. T. , Selection and Performance of Vibration Tests, The Shock and Vibration Information Center,

SVM-8 (1971).

34. Fackler, W. C, Equivalence Techniques for Vibration Testing, The Shock

and Vibration Information Center, SVM-9 (1972).

35. Zimmerman, D., Johns Hopkins Applied Physics Laboratory, private communication.

36. Information on Military Standards can be obtained from the Naval Publications and Forms Center, 5801 Tabor Avenue, Philadelphia, PA 19120.

37. Ebel, G. , Singer Corporation, comments made at the Moisture Measurement

Workshop, NBS, November 5-6, 1980.

38. GIDEP Alerts may be obtained from the Government-Industry Data Exchange

Program, Corona, CA 91720,

39. Thomas, R. W. , Moisture, Myths and Microcircui ts , IEEE Trans, PartSj

Hybrids, and Packaging PHP-12 , 167-172 (September 1976).

40. See, for instance, Espe, w. , Materials of High Vacuum Technology, Vol.

1-3 (Pergamon Press, New York, 1966).

41. Gulati, S. T. , and Hagy, H. F. , Expansion Measurement Using Short Cylindrical Seal: Theory and Experiment, Thermal Earpansion, E. D. Peggs,

Ed., Vol. 6 (Plenum Press, New York, 1978).

42. Information on ASTM Standards may be obtained from the American Society

for Testing and Materials, 1916 Race Street, Philadelphia, PA 19103.

43. McCormick, J., and Zakraysek, L. , A Metallographic Test for Glass-toMetal Seal Quality, 17th Annual Proc, Reliability Physics, San Francisco, California, April 24-26, 1979, pp. 44-50.

44. See, for instance, Griffel, w. , Handbook for Formulas for Stress and

Strain (Frederick Ungar Publishing Co., New York, 1967).

45. This and other ISHM publications are obtainable from the International

Society for Hybrid Microelectronics, P.O. Box 3255, Montgomery, AL

361 09.

46. Roark, R. J., and Young, w, C, Formulas for Stress and Strain

(McGraw-Hill, New York, 1975).

^ U.S. GOVERNMENT PRINTING OFFICE: 1982-360-997/2114

73
NBS-n4A (REV. 2-BC) U.S. DEPT. OF COMM. 1. PUBLICATION OR 2. Performing Organ. Report No. 3. Publication Date BIBLIOGRAPHIC DATA

REPORT NO.

May 1982

SHEET (See instructions) NBS SP 400-70

4. TITLE AND SUBTITLE

Semiaonductor Measurement Technology : The Use of Acoustic Emission to Determine

the Integrity of Large Kovar Glass-Sealed Microelectronic Packages

5. AUTHOR(S)

G. G. Harman

6. PERFORMING ORGANIZATION {if joint or other thon NBS. see instructions) 7. Contract/Grant No. NATIONAL BUREAU OF STANDARDS P.O. N0016380MP0001

DEPARTMENT OF COMMERCE 8. Type of Report & Period Covered

WASHINGTON, D.C. 20234

Final

9. SPONSORING ORGANIZATION NAME AND COMPLETE ADDRESS (Street. City. State, ZIP)

Naval Avionics Center

Indianapolis, IN 46218

10. SUPPLEMENTARY NOTES

Library of Congress Catalog Card Number: 82-600528

Document describes a computer program; SF-185, FlPS Software Summary, is attached.

11. ABSTRACT (A 200-word or less factual summary of most significant information. If document includes a significant

bi bliography or literature survey, mention it here] The general objective of this research was to develop tests to determine the

integrity of large hybrid packages under various thermal and mechanical stresses

that may be encountered during assembly, during installation in systems, or in

operation. Several measurement techniques were investigated, but emphasis was placed on acoustic-emission test procedures. The accomplishments were: (1) The

effects of avionics environmental vibration on the seals of hybrid packages mounted

on printed-circui t (PC) boards were determined. A major conclusion of this section

was that lead fatigue failure occurs before seal damage on packages from high

quality lots. (2) A small acoustic-emission detector was developed that is

sensitive to surface waves, but relatively insensitive to vibration induced cable

noise. (3) A high-temperature (125°C) open-package helium leak test method was successfully developed to observe marginal seal damage. (4) An acoustic-emission

test for inspection of hybrid packages during high-temperature thermal shock was developed. (5) A study of possible damage to seals during thermocompression and

thermosonic bonding, during lead forming, and during other assembly operations was carried out. A general conclusion of this study is that the glass-metal seals in

packages from known high quality lots are very reliable even when subjected to high

stresses. However, the seals from packages "screened as good" from reject or poor

quality lots are subject to hermetic failure under moderate stresses. There is

little correlation between visual inspection failures of glass seals and their

hermeticity.

12. KEY WORDS (Six to twelve entries; alphabetical order; capitalize only proper names; and separate key words by semicolon s) acoustic emission; hermeticity; hybrid microelectronics; hybrid packages; microelectronic packaging; thermal shock; vibration

13. AVAILABILITY

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r~] Order From Superintendent of Documents, U.S. Government Printing Office, Washington, D.C. 20402.

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80

15. Price

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