533

Fracture Properties of Molding Compound Materials for IC Plastic Packaging

John Sauber, Lidia Lee, Shih Hsu, and Trirat Hongsmatip

Abstruct- The technique of linear elastic fracture mechanics (LEFM) was employed to characterize the fracture toughness different molding compound materials. The effect of fast thermal loading rate, as in a wave soldering condition, was studied by performing the fracture toughness tests at different mechanical loading speeds. The effects of storage conditions and accelerated testing environments were studied by varying the test tempera- tures from liquid nitrogen temperature to 150°C.

Models were built of 208 I/O PQFP devices with cracks in the molding compound at the corner of the die pad. These models were solved to evaluate the effect of CTE mismatches, initial flaw sizes, and die pad delamination on molding compound stress intensity factors. Finite element results were then compared with crack growth measurements from PQFP packages that had been subjected to accelerated thermal cycling.

I. INTRODUCTION

LASTIC packages have been used extensively by the P electronics industry due to the cost-saving advantage. Demand for replacing a relatively expensive ceramic package with a thermally and electrically enhanced plastic package for high-power chips has been increasing. However, the resulting package construction become more complex, such as in a thermally enhanced plastic package containing a metal heat spreader where sharp anchoring comers may be buried. A successful implementation therefore requires a good knowl- edge of the mechanical performance of the materials in various process and operational environments.

Most mechanical analyses of plastic packages employ con- ventional continuum mechanics to determine the stress field acting at various parts of the packages due to thermal ex- pansion differentials. Interfacial adhesion has been assumed to be perfect in many cases. Unfortunately, local defects are often present in as-received IC packages, in the form of local delamination or preexisting cracks at high stress concentration areas such as the edges of the chip and the die bonding pad. These areas are typically very sharp, with a small radius of curvature. Continuum mechanics cannot be used in this case due to the stress singularity at the crack tip. If the finite element method is applied, the computed maximum stress value at these sharp comers is dependent on the element size around the point and no definite values can be obtained.

The technique of fracture mechanics must be used to predict stress field around the crack tip. However, there are a limited

Manuscript received February, 1994; revised July 18, 1994. This paper was presented at the 44th Electronic Components and Technology Conference, Washington, DC, May 1 4 , 1994.

The authors are with Digital Equipment Corporation, Hudson, MA 01749 USA.

IEEE Log Number 94059 12.

number of examples that show the application of this technique in predicting mechanical failure of electronic packages [l], [2]. One such example has been demonstrated by Nishimura er al., who used linear elastic fracture mechanics (LEFM) to examine fatigue crack propagation of a plastic package under repeated thermal loading conditions [ 11.

This study demonstrates the use of an LEFM technique in characterizing the fracture properties of three molding compound materials that have been selected for applications such as PQFP and TSOP products. In addition, finite element models were built to evaluate the effect of CTE mismatches, initial flaw sizes, and die pad delamination on molding com- pound fracture properties. A coupling of the fracture toughness values obtained from testing with results from finite element models allowed a failure criterion to be established.

11. LINEAR ELASTIC FRACTURE MECHANICS Crack propagation in a brittle polymer has been explained

using several fracture theories, which fall generally into two groups: energy criteria and stress criteria theories [2]-[4]. For the latter, a stress intensity factor is a measure of the magnitude of the crack tip stress field. It depends on the gross section stress o and the square root of the half crack length c. Thus the stress intensity factor K incorporates both geometrical terms and the stress level. For various geometries with cracks of length 2c,

K = Y &

where Y = geometrical factor that depends on geometry and loading conditions.

If the crack extends at nominal stresses significantly less than the yield stress, and if the K at fracture is found to be a constant for various specimen geometries, this value is denoted as K,. The value of K, provides a single parameter fracture criterion, the critical stress intensity factor at which the crack will extend rapidly.

A material with a high K, has a greater resistance to crack propagation and therefore a greater mechanical reliability. The fracture toughness value of a molding compound material depends on the chemical makeup of its resin system and filler content, shape, and size. In addition, the values of K, could vary with temperature and rate, as different molecular energy absorption mechanisms may be active in different environments.

1070-9886/94$04.00 0 1994 IEEE

534 IEEE TRANSACTIONS ON COMPONENTS, PACKAGING, AND MANUFACTURING TECHNOLO(;Y-PART A, VOL. 17, NO. 4. DECEMBER 1994

TABLE I MOLDING COMFOUND bhT'EKIAL PROPERTIES

TABLE 11 EFFFCT OF TFMPERAI U N OU FRACTURE TO1 h H N E 6 S

Temp.

-196 1.76 2.49 2.30 3.73

85 1.93 1.98 4.05 I 155 1 0.12 1 1.90 1 0.68 I

111. TESTING TECHNIQUE

In this study, K , values of various molding compound materials were characterized using a single-edge notch bending technique. A very sharp crack tip was introduced in the V- shaped notch of each sample using a heavy duty utility blade, simulating a natural crack tip.

For a sample with crack length U, width W, thickness B, and distance between supports S , the K , value is calculated based on the critical applied load (1;) using the ASTM standard method [ 5 ] , [6]:

where

7 / 2 - 37.6 (G) + Y8.q 6) 9'2

The effect of various thermal loading conditions, i.e., slow, vs. fast, on K, value5 were studied by performing the fracture toughness tests at different mechanical loading speeds: 2.54 and 25.4 "/minute. Effects of temperature related to storage conditions and accelerated testing environments were studied by varying the test temperatures from liquid nitrogen temper- ature (-196°C) to 155°C. Although -196°C is unrealistically low, the K, value at this temperature could be used as the most conservative value. Furthermore, evidence suggests that the h', value determined at this temperature could be close to the K, value under a normal subambient storage condition, given the fact that K,. generally levels off at subambient temperatures. The test was done by immersing; the samples in liquid nitrogen for at least 15 min until the nitrogen stopped boiling. The samples were then tested in the Instron immediately after they were removed from the liquid nitrogen.

Iv. MATERIALS

Three different molding compounds (A, B, and C) were tested (see Table I) Materials .4 and B were standard epoxy cresol novolac (ECN) based PQFP molding materials, while material C was a low-stress, biphenyl-based molding material. Characteristics of a typical biphenyl-bascd molding compound can be found in [7].

The single-edge notch bending samples made of materials B and C were individually molded through a transfer molding process. Samples from material A were cut from a molding sheet, as suggested by the surface finishes.

V. TEsr REX'

A. Effect o j Tcrriperrrtiue on K,. The stress intensity factor near tt < tip under a low-

temperature storage condition is exp( o be different from that under a solder reflow condition o the sensitivity of mechanical strength and modulus to Iperature. To model failures under difierent environmental conditions, K,. values of the three molding compound materials at different temper- atures were obtained as shown in Table 11. The tests were performed at a rate of 2.54 nim/min. The K,. values reported were averages of four samples. Standard deviations of the values were less than 0.2 MN/m"/' in all cases.

In general, the K, values decreased as the temperature de- creased from room temperature to a subambient condition for the materials. This implies that cracks tend to grow more easily as the packages were cooled to a subambient temperature. This was likely due to an inability of the molecular chains to undergo deformation at the low temperatures.

Within the temperature range of - 196°C to 155"C, the K, values for materials A and B are quite close to each other, suggesting that materials A and B have comparable fracture resistance under a normal temperature cycling conditions. However, the K,. value for material A was significantly lower, at 155°C. This is possibly due to the fact the test temperature exceeded the T, of the material. At the solder reflow tem- perature (24OoC), it IS possible that fracture resistance of the two materials would become comparable. Material C showed noticeably higher K , values within the -196°C to 155°C range. This is suggested to be due to the low-stress chemical formulation of the biphenyl molding compound. In addition, a higher filler loading in the biphenyl molding compound could have contributed to the higher K , values observed, as it has been demonstrated that the K( value of a molding compound increases with its filler loading and particle size [SI.

Except at 155"C, all of the samples exhibited brittle fracture, i.e., the crack propagated rapidly when a critical load was reached. At 155"C, however, the crack grew slowly as the critical load had passed. It i s suspected that the slow crack

SAUREK et U / . : MOLDING COMPOUND MATERIALS FOK IC PLASTIC PACKAGING

Loading Temp. Rate ("C)

("/mi n .)

535

K, (MN/m"L) for Moking Compound

A [ B [ C

Fig. I . Test sample crack morphology.

155 -196 20 85

254

Fig. 2. refow.

PQFP crack morphology after moisture saturation and vapor phase

0.12 1.90 0.68 1.59 1.03 1.17

1.51 3.36

growth was governed by the viscoelastic properties of the material and the crack tip was proceeded by a sizable yield zone. The use of linear elastic fracture mechanics might not have been valid in this case.

Adhesion between the filler particles and the resin matrix appeared to be good in all the cases, as suggested by the fact that crack propagated through the resin matrix, leaving filler particles unexposed. A typical fracture morphology ob- tained is shown in Fig. I . Although the sample shown was tested at room temperature, fracture morphology did not vary significantly with the test temperatures or the loading rates.

This is in contrast with a previous observation, where the crack was found to propagate along the residfiller interface in a PQFP sample that had gone through a saturated moisture test followed by a vapor phase reflow process, as shown in Fig. 2.

It is thought that the moisture had an adverse effect on the interfacial adhesion between the resin and the filler particles, thus giving rise to a diffcrent crack propagation mode.

8. Efiect of Thermal Loading Kate

To simulate the loading from accelerated thermal cycling with a mechanical loading test. a correlation must be es-

TABLE 111 EFFECT OF LOADlUG RAT^ ON FKACTUKI: ToLGti\tSS

I -p 1 1.76 1 1.98 1 22; 2.49 2.30 1.93 1.98 4.05

2.54

tablished between the crack opening displacement induced by mechanical loading and the amount of crack opening due to expansion differentials during thermal cycling. Based on an FEM calculation of the notched specimen, a thermal cycling condition of -55°C to 125°C in I O min (0.3"C/s) will cause a 0.013 I d s crack opening diqplacement rate. The mechanical testing speed corresponding to this rate of crack opening displacement is 3.5 ,um/min or 0.042 i d s . As another example, if the ramp rate during vapor phase reflow is 4"C/s, this translates to a mechanical testing speed of 0.03 "/min.

The effect of mechanical loading rate on fracture toughness was determined. Results are shown in Table 111. The standard deviations for the values reported were all less than 0.18 MN/m:'/*.

A clear effect of mechanical loading rate on K(. was evident. Kc decreases with an increase in the mechanical loading rate or, in turn, an increase in the thermal loading rate. However, the slowest mechanical testing rate used was higher than the calculated rate for a vapor phase reflow condition.

VI. FINITE-ELEMENT MODELING OF PQFP DEVICES

Finite element models of 208 I/O PQFP devices with cracks in the molding compound at the corner of the die pad were used to evaluate the effect of CTE mismatches, initial flaw sizes, and die pad delamination on molding compound stress intensity factors.

A. Modeling Crack Tips

Before building the PQFP models, a preliminary study was conducted to determine the best method for modeling a crack tip. Models of the notched test specimens were built using four different crack tip configurations. These were U-notch. V-notch, keyhole, and singular element. The models were solved using the ABAQUS' finite element code to evaluate the J-integral at the crack tip. The J-integral is identical to the strain energy release rate (G) and the crack driving force 191. For a linear plane strain problem, the relationship between the

' ABAQUS is a registered trademark of Hibbett, Karlason & Sorensen, Inc.

536 IEkF TRANSACTION9 ON COMPONENTS, PACKAGING, AND MANUFACTURING TECHNOLOGY-PART A, VOL 17, NO 4, DECEMBER 1994

. <-e* r-r ,..

die< die attach b

die pad

crack

...AL..

Fig. 3. Typical model of PQFP device (close-up) view of corner of die pad.

J-integral and the stress intensily factor K is [IO], I1 I ]

where E is the elastic modulus and U i? Poisson's ratio. All of the configurations except the V-notch produced results within 2% of the closed-form solution. Since the keyhole was the easiest to construct, it was used for all subsequent crack tip modeling.

An additional study wax conducted to determine how small the crack tip radius must be to produce accurate results. A series of models werc solved with crack tip radii ranging from 0.001 mm through 1 mm. It wiis found that the same answer was obtained for any radius of 0.02 mm or less, so this was the value used in thc PQFP models.

B. PQFP Device Modeling

Models of diagonal cross sections of PQFP devices were built using higher order (%node) plane strain elements. A plane strain element formulation was used because it includes the effect of Poisson's ratio and has been shown to produce a reasonably good approximation1 of the 3-D stress state within plastic packages [12]. Due tal symmetry, only half of the package was modeled. These models included a die, die pad, die attach, and molding compound, as shown in Fig. 3 .

Models also included a crack that extended downward at a 15" angle from the comer of the die pad. This crack location is based on microsections of devices after thermal cycling and also concurs with observations made by other researchers [ 11 .

The loading applied to the models was cooling from a zero stress state at 175°C to -55"C, which is the lowest temperature reached during qua.lification testing. Since a cycle length of 20 min is used for the test, it is assumed that thermal gradients within the package are negligible and that a steady state solution will produce acceptable results.

IO 15 20 25

MOLDING C O M P O U N D C T E [ P P M / C I Fig. 4. with 0. I3 mm crack and delamination type B.

Stress intensity factor versus molding compound CTE (below Tq)

Except where otherwise noted, the material properties shown in Table IV were used. The molding compound Tg used was 150°C with a CTE above Tg of 63 ppm/"C.

C. Effect of Delamination

While delamination of the bottom of the die pad is typical for devices with Alloy 42 die pads, microsections of devices with copper alloy die pads showed several different types of delamination after thermal cycling. Models were built with sliding surfaces at the interfaces to determine the impact of the various types of delamination on the stress intensity factor at the tip of a 0.13 mm long crack. Table IV shows a summary of these results. The column labeled "Probability" indicates the likelihood of seeing a particular type of delamination after 100 accelerated thermal cycles.

The results show that an understanding of which surfaces will delaminate is essential for accurately establishing a failure criteria. It was decided that delamination type B would be used as the typical loading condition.

SAUBER er al.: MOLDING COMPOUND MATERIALS FOR IC PLASTIC PACKAGING

Material

Die

537

E CTE (GPa) (ppml"C) 95.7 2.7

TABLE IV MATERIAL PROPERTIES FOR FEA MODELS

Die Attach Alloy 42 Die Pad & Leadframe Cu Alloy Die Pad & Leadframe

8.6 65 146.1 5.9 118.5 17

Molding Compound 11.8 16

TABLE V

(0.13 mm CRACK, COPPER ALLOY DIE PAD) IMPACT OF VARIOUS TYPES OF DELAMINATION ON STRESS INTENSITY FACTORS

Type of Delamination Prob- K, ability

RAN/~I~'~

A B -

- C

- D

No delamination 20% of die pad bottom Die pad side Die pad top (to edge of die) Side of die 20% of die attach Same as "B" except: 100% of die pad bottom No die attach delam. Same as "B" except: 100% of die pad bottom

rare 0.14 most 0.72

common

occasional 0.99

occasional 1.00

E

D. The Effect of CTE Mismatches

The CTE of the various molding compounds was measured using a thermomechanical analysis machine. While the mean values measured were fairly close to the vendors' published values, there was a considerable amount of scatter in the data (standard deviation between 1 and 1.7 ppm "C for the range below Tg). Because of this, a series of models were solved to evaluate the effect that molding compound CTE has on the stress intensity factor at the tip of a 0.013 mm crack. The delamination used for these models is type B, as described in Table V.

It can be seen that using a lower CTE molding compound reduces the stress intensity factor when either die pad material is used. This is not surprising since the most significant expansion differential is with the die that has a low CTE. The Alloy 42 die pad works "with" the die, which increases the molding compound stress, while the copper alloy die pad works "against" the die, which reduces the molding compound stress.

Unfortunately, the most common way to reduce the CTE of a molding compound is to increase the amount of silica used for filler. This usually increases the modulus, which in turn can increase the stress. An increased filler content can also result in a lower fracture toughness. Therefore, these factors

! rare 1 Same as "B" except: 100% of die pad bottom 100% of die attach

must be considered before recommending a change to a low CTE molding compound.

It is worth noting that the failure mechanism is the same regardless of whether the CTE of the die pad is higher or lower than that of the molding compound. Figs. 5 and 6 shows that in one case the die pushes the molding compound sideways, and in the other case the pad pushes the molding compound sideways, but in either case the net result is a Type I crack opening displacement, which occurs in the molding compound at the bottom comer of the pad.

This is not a new observation [13], but it bears repeating since it shows that the thermal loading inside a plastic package is very similar to the loading that the notched specimens are subjected to during fracture toughness testing, i.e., they both produce a Type I crack opening displacement.

E. The Effect of Crack Length

Since the stress intensity factor at the tip of a crack depends on the length of the crack, a study was undertaken to evaluate the changes that occur as the crack length increases. A series of models were solved including an Alloy 42 die pad and sliding surfaces to simulate delamination of the interface between the bottomhide of the die pad and the molding compound. It has been observed that this sort of delamination is fairly common and often occurs after only a few thermal cycles [ 131. The results are shown in Fig. 7.

For the Alloy 42 die pad, initially the stress intensity factor decreases as the crack tip gets away from the stress concentration caused by the comer of the die pad pushing on the molding compound (see Fig. 5) . As the crack continues to grow, the stress intensity factor increases since the distance to the outside of the package is decreasing and there is less material left to restrain the load.

This behavior of K increasing as the crack length increases is the same as that exhibited by the notched specimens (see (2)). It is also the only factor that influences the effect of crack length when a copper alloy die pad is used. Since a copper alloy die pad does not push directly against the molding compound (see Fig. 6), the initial decrease in the stress intensity factor with increasing crack length does not occur.

VII. PQFP TESTING A set of 32 samples of 208 U 0 PQFP devices from two

different vendors (X and Y) were subjected to an accelerated thermal cycling plan. The devices had copper alloy die pads and lead frames, a 10.3 mm square die. and used molding

538 lEEE TR4NSACTIONZ ON COMPONENTS, PACKAGING. AND MANUFACTURING TECHNOLOGY-PART A, VOL. 17, NO. 4, DECEMBER 1994

Fig. 5 . Deformed shape when CTE of die pad (o = G ppm/OC) is lower than CTE of molding compound (01 = 20 ppm/"C).

compound A. The thermal cycle was from -55°C to 125"C, with a 20-min total cycle length. After 0, 100, 200, 500, 1000, 2000, and 4000 cycles, four samples were removed from the test and subjected to C-SCAN to evaluate the amount of delamination of the die and die pad. Microsectioning was also done to measure the length of cracks in the molding compound.

The C-SCAN results showed that the amount of die pad delamination increased only slightly as a result of thermal cycling. Delamination was initially only 5%) to 10% of the pad area, and then gradually increased to - 20% at 4000 cycles. There were also a few cases of 20%# to 60%: delamination, which were randomly scattered and did not correlate with the number of cycles.

Delamination of the die attach occurred in devices from both vendors but began much earlier (200 vs. 2000 cycles) in the devices from vendor X. This appeared to be caused by an insufficient amount of die attach material being applied. This gap between the die and pad was filled with molding compound during encapsulation. The resulting stress concen- tration at the corner of the die apparently causes delamination to begin earlier and grow faster than it did in devices from vendor Y (see Fig. 8).

A least squared fit to the data shows that the crack growth rate is 2.28-5 mmkycle for vemdor X and 1.5E-5 "/cycle

for vendor Y. If the cracks continued to grow at the same rate they would reach the outside of the package (1.22 mm) in 52 000 and 78 000 cycles for vendors X and Y, respectively. These estimates are unconservative since the stress intensity factor increases slightly as the crack length increases (see Fig. 7). which would result in an increased rate of crack growth. It is also possible that this increase could result in K becoming greater than K ~ C . , which would cause the crack to become unstable and result i n immediate failure of the device.

The difference between the two vendors can probably be attributed to the previously mentioned differences in the die atlach. Since the insufficient die attach led to delamination which began earlier and grow faster, and since increased delamination causes an increase in the stress intensity factor (Table V) it makes sense that devices from vendor X would have a higher rate of crack growth.

Devices were sectioned to show the amount of crack growth that occurred at the corner of the die pad. These results showed that crack length increases as a result of thermal cycling, as shown in Fig. 9.

VIII. MODELING RESULTS VERSUS TEST RESULTS

The stress intensity factors K calculated for the actual PQFP devices that were tested are shown in Table VI. These results

SAUBER et al.: MOLDING COMPOUND MATERIALS FOR IC PLASTIC PACKAGING

Fig. 6. Deformed shape with CTE of die pad (0 = 17 ppml°C) is higher than CTE of molding compound (nl = 12 ppmPCj.

0 . . 2 0 . 4 0 .60 .BO 1 .o 1 . 2

C R R C K L E N G T H [ M M I Stress intensity factor vs. crack length with delaminated die pad. Fig. 7.

are based on molding compound A (used for both types of PQFP’s), delamination type B (from Table V), a 0.13-nim crack length, and a standard deviation (T on the molding compound CTE of I ppm/”C (based on our test results).

By interpolating the fracture toughness data for molding compound A, we can determine that at -55°C its Klc value is approximately 2.24 MN/m”’ (see Table 11). So, for

Fig. 8. (vendor X).

Delamination resulting from an insufficient amount of die attach

nominal conditions the factor of safety is large (0.722 vs. 2.24 MN/m”’). However, assuming worst-case conditions shows a much lower margin of safety.

Delamination conditions C and D are uncommon but do occur, and either condition will result in a 39% increase in

540 IEEE TRANSACTIONS ON COMPONENTS, PACKAGING, AND MANUFACTURING TECHNOLOGY-PART A, VOL. 17, NO. 4, DECEMBER 1994

0 1000 2000 3000 4000 5000

NUMBER O F THERMAL C Y C L E S Fig. 9. Crack length versus number of‘ thermal cycles.

TABLE. VI MODELING RESULTS FOR THE WFP DEVICES THAT WERE TESTED

Molding Compound CTE

Nominal - 30 = 13 0.535 INominal = 16 1 0.722 I Nominal + 30 = 19 1 0.909

the value of K (see Table V). If the increased delamination is combined with the Nominal f3a molding compound CTE ( 1 9 ppm/”C), the stress intensity factor is increased to 1.26 MN/m3/*, which will result in an increased crack growth rate.

Furthermore, as the crack length increases, so does the stress intensity factor (see Fig. 5). As the crack approaches the outside of the package, the K value is increased by another 23% to 1.55 MN/m”’. In addition, we know that the standard deviation on the fracture toughness test results is 0.2 MN/m3/’, so using a Nominal -3a value reduces the fracture toughness to 1.64 MN/m312.

Fortunately, this worst-case loughness value is still greater than the worst-case stress intensity factor, and it is extremely improbable that all of these conditions will occur simultane- ously (odds are less than I in IO‘)), so it is unlikely that rapid crack growthlimmediate failure will occur. While slow crack growth will eventually result in device failure, it will require many thousands of cycles before the crack reaches the outside of the package. So even in the worst case it appears that these PQFP devices have an acceptable factor of safety.

IX. CONCLUSIONS

To summarize the major results of this study, the following

The technique of linear elastic fracture mechanics was used to determine the effects of temperature and loading rate on the fracture toughness of molding compounds used for IC encapsulation. Finite element models of PQFP devices with cracks in the molding compound itt the comer of the die pad

points can be made.

were used to evaluate the effect of die pad delamination, CTE mismatches, and crack length on molding compound stress intensity factors. Delamination has a significant effect on the stress in- tensity factor. An understanding of which surfaces will delaminate is essential for accurately establishing a failure criteria. Using a lower CTE molding compound reduces the stress intensity factor when either die pad material is used. The use of a copper alloy die pad results in lower stresses than with an Alloy 42 die pad. Regardless of whether the CTE of the die pad is higher or lower than that of the molding compound, the net result is a Type 1 crack opening displacement at the bottom comer of the die pad. With an Alloy 42 die pad, the stress intensity factor at th crack tip decreases slightly as the crack grows, and then increases again as the crack approaches the outside of the package. With a copper alloy die pad the stress intensity factor increases steadily as the crack grows. Based on a comparison of PQFP modeling results with fracture toughness testing results, PQFP devices have a large factor of safety under nominal conditions and an acceptable factor of safety under worst-case conditions.

ACKNOWLEDGMENT

The authors thank H. Gordon, J . Ide, G. Morand, J. Oparowski, T. McCoy, and H. Nguyen, whose efforts made this work possible.

REFERENCES

[ 11 A. Nishimura. A. Tatemichi, H. Miura, and T. Sakamoto, “Life es- timation for IC plastic packages undcr temperature cycling based on fracture mechanics,” IEEE Truns. Comp., Hybrids, Munuf: Techno/., vol. CHMT-12. no. 4, p. 637, Dec. 1987.

[Zl A. Dasgupta and J. M. Hu, “Fracturc mechanism models for brittle fracture,” /EEE Truris. Reliubiliry. vol. 41, no. 3, p. 328, 1992.

[31 D. Broek, Elenienrcry Enxinerrin,q Fructure Mechunics. Alphen aan den Rijn, The Netherlands: Sijthoff & Noordhoff, 1978.

[4] J. F. Knott, Fundumenfuls of Frncfurc, Mechunics. London: Butter- worths, 1973.

[ S ] “Standard test method for plane-$train fracture toughness of metallic materials,” ASTM E399

[6] J . E. Srawley, and W. F. Browm. “Fracturc toughncss testing methods,” ASTM STD 381, pp. 133-14.5. 196.5.

[71 N. Mogi and H. Yasuda, “llevclopment of high-reliahility epoxy mold- ing compounds for surface-mount devices,” in IEEE Proc. 42nd Elecrro. Comp. & Tech. Cor75. 1992, p. 1023.

[8] M. Yamaguchi, Y. Nakamura, M. Okubo, and T. Matsumoto, “Strength and fracture toughnexs of cpoxy resin tilled with silica particles,” Nitto Denko Tech. Rep.. Jan. 1091, pp. 74.

[ V I H. D. Hibbett, in “l’racticnl and computational aspects of fracture me- chanics,” Cour,ve Nofc,s from A d i x “ d Tr,pic..s in Fruc.fure. Pawtucket, RI: Hibbett. Karlsson & Sorenscn, 1989, ch. 4, p. 50.

I I O ] -, “Linear elastic J-integral calculations for tension and bend- ing specimens,” in C o u r . ~ Noie.v f r o n ~ Advanced 7opies in Fracture. Pawtucket, RI: Hihhett, Korlsson & Sorenscn, 1989, ch. 4, p. 103.

[ I I ] M . Kutr., Mec.htrnicri1 E ~ R I I I P ~ J ~ U ’ Hundhook. New York: Wiley, 1986,

[I21 G. Kelly, C. Lyden. C. O’Mathuna, and J . S. Campbell. “Investigation of thermo-mechanically induced stress in a PQFP 160 using finite element technique\.” in /I:EE Pruc.. 42nd Electro. Comp. Con$. 1992, pp. 467472.

p. 264.

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1131 A. Nishimura and S. Kawai, “Effect of leadframe material on plastic- encapsulated IC package cracking under temperature cycling,” in IEEE Pmc. 39th E l ~ t r ~ . Comp. Con/:. 19x9, pp. 524530.

John Sauher received the B.S.M.E. degree from Northeastern University in 1983.

He is with Digital Equipment Corporation’s Semiconductor Operations group. For the last eleven years hc has been primarily involvcd in developing finite element methods for predicting the reliability of various types of electronic component\ and assemblies. This includes developing a fracture mechanics approach for evaluating plastic packaging options, a method for predicting fatigue lifetimes for surface mount solder joints, and

methods for evaluating PWB reliability in random vibration environments. Prior to joining Digital in 1987, he was involved in the design and analysis of high-density interconnects for both Texas Instruments and Midland Ross Corporation.

Lidia Lee received the Ph.D. degree from Mas\- achusetts Institute of Technology, Cambridge. in polymer science & engineering in 19x4. Her study focused on polymer based composites and fracturc of polymers. She also received the M.S. degree from Polytechnic Institute.

She has been working on materials applications in the areas of semiconductor packaging and intercon- nect technology at Digital Equipment Corporation for the last eight years. Her recent involvement has been in the area of plastic packaging. Prior to this

she worked for Army Materials Research Lab, Watertown,-MA

Shih Hsu, photograph and biography not available at time of publication

Trirat Hongsmatip, photograph and biography not available at time of publication.