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IEEE TRANSACTIONS ON COMPONENTS AND PACKAGING TECHNOLOGIES, VOL. 28, NO. 4, DECEMBER 2005

Swelling and Time-Dependent Subcritical Debonding of Underfill During Temperature–Humidity Aging of Flip Chip Packages E. H. Wong, Ranjan Rajoo, Thiam Beng Lim, Member, IEEE, and Yui-Wing Mai

Abstract—Stress corrosion cracking and the associated time-dependent subcritical debonding of the underfill–die interface is believed to be responsible for a large disparity between the reliabilities of flip chip–underfill packages subjected to two 3500 h) and different aging conditions: 1) 85 C/85% RH ( 2) 121 C/100% RH ( 100 h). The ratio of the transverse stress at the bump (Sy) to the shear stress at the die edge (Sxy) has been found to correlate well with the two characteristic autoclave failure signatures: random local failure and die edge delamination. The first failure mode is promoted by a high Sy–Sxy ratio. A practical finding from this study is that raising the curing temperature of the underfill offers a simple and effective means of enhancing the autoclave performance of the flip chip–underfill package. Index Terms—Debonding, hygroscopic swelling, swelling, temperature–humidity (T–H) aging.

I. INTRODUCTION

Fig. 1. Hygroscopic swelling induced tensile failure in autoclave testing of a flip chip PBGA package. TABLE I AUTOCLAVE LIFE OF PACKAGES VERSUS HYGROSCOPIC STRAIN OF UNDERFILL MATERIALS

M

ICROSCOPICALLY, moisture can be present in a polymer structure in two basic forms: 1) physically resident in molecular-sized holes (free volume) and 2) chemically attached to the hydrogen bonding sites along the polymer chains. The availability of holes depends on the polymer microstructure, morphology, and crosslink density, whereas chemical attachment depends on the presence of hydrogen bonding sites. The unbound moisture that resides in the free volume does not lead to dimensional changes of the polymer. Chemically bonded water molecules, on the other hand, disrupt the interchain hydrogen bonding and lead to swelling and plasticization of the polymer [1], [2]. Differential swelling occurs between polymeric and nonpolymeric materials as well as among the various polymeric materials constituting the electronic packages. This differential swelling induces hygroscopic strain and stress in the electronic package that act in tandem with thermally induced strain and stress. In many instances, the magnitude of the moisture-induced stress may even be higher than the thermally induced stress [3], [4].

Manuscript received May 30, 2003; revised March 15, 2005. This work was recommended by Associate Editor C. P. Wong upon evaluation of the reviewers’ comments. E. H. Wong is with the Institute of Microelectronics, Singapore 117685, and also with the University of Sydney, Sydney, NSW 2006 Australia (e-mail: [email protected]). R. Rajoo and T. B. Lim are with the Institute of Microelectronics, Singapore 117685. Y.-W. Mai is with the University of Sydney, Sydney, NSW 2006 Australia. Digital Object Identifier 10.1109/TCAPT.2005.850526

Flip chip assemblies using underfill are known to be susceptible to autoclave failure [3]–[7]. The introduction of underfill, intended as reinforcement, may even degrade the autoclave life of a flip chip assembly [4], [6]. The mechanism involved is hygroscopic swelling of the underfill, which induces transverse tension that cracks the solder interconnection, as illustrated in Fig. 1. Table I shows the results from earlier experiments [4] that demonstrate excellent correlation between the autoclave life of flip assemblies and hygroscopic strain, for six underfill materials. Despite the excellent correlation, two unanswered questions remain, which are the focus of this study: 1) why is there a large disparity between the autoclave life and the 85 C/85% RH life of a flip chip–underfill package? and 2) why are there two characteristic failure modes in the autoclave failure? To answer these questions, four representative underfill materials out of the previous six studied were further evaluated in the autoclave (121 C/100% RH) and 85 C/85% RH aging experiments. The moisture-related properties—saturated concentration, diffusivity, and hygroscopic swelling—of these underfill

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WONG et al.: SWELLING AND TIME-DEPENDENT SUBCRITICAL DEBONDING OF UNDERFILL DURING TEMPERATURE–HUMIDITY AGING

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TABLE II AUTOCLAVE AND 85 C/85% RH STORAGE PERFORMANCE

Fig. 2. Schematic illustration of technique for characterizing hygroscopic swelling.

materials were characterized in detail. Finite-element analyses were performed to provide insights into the results. II. DESCRIPTION OF EXPERIMENTS A. Temperature–Humidity (T–H) Aging A flip chip plastic ball grid array (PBGA) package with a 10 10 mm daisy-chained chip and 27 27 mm substrate was used as the test vehicle. The packages were assembled with four commercial underfill materials from different sources. The coding for these underfill materials is indicated in Table I. For each underfill material, 15 package assemblies were subjected to 85 C/85% RH aging while 12 were subjected to 121 C/100% RH aging. Electrical connectivity was monitored at intervals during the aging experiments. Electrical open was considered as a failure. The packages were acoustically inspected prior to the aging and at the time of electrical failure. B. Hygroscopic Swelling Characterization The technique for characterising the hygroscopic swelling of a polymeric material using a combination of thermomechanical analysis (TMA) and thermogravimetric analysis (TGA) has been presented in earlier publications [3], [4] and is illustrated in Fig. 2. The underfill materials as well as the organic substrate used in this work were prepared in specimen dimensions of 4 4 1 mm. When a linear relation exists between the hygroscopic strain and the moisture concentration, the linear constant is defined as the coefficient of moisture expansion (CME) [3], [4], [8]. Mathematically, hygroscopic swelling strain can be defined as CME

terized at 85 C/85% RH using square specimens of dimensions 50 50 1 mm, in accordance to the standard procedure [9]. Similar characterizations were not possible at the autoclave conditions (121 C/100% RH) because of the depressurization sequence at the end of the conditioning process that results in significant loss of moisture. Instead, Csat at 121 C/100% RH was projected from the 85 C/85% RH data using the following assumptions: 1) Moisture absorption obeys Henry’s law, that is, Csat is linearly proportional to the environmental humidity. In reality, the linear relation is valid only for low moisture concentrations. A better description of the relation is given by (2) [10] Csat

RH

where and are constants and is greater than one for most polymers at relative humidity above 60% [11]. Hence, the linear assumption will lead to underestimation of Csat at 121 C/100% RH. 2) Csat is independent of temperature, which is a good engineering approximation [12], [13]. The diffusivities of the underfill materials at 85 C and 121 C were extracted from the TGA desorption measurement. from a Describing the total mass of the moisture desorbed rectangular element of a specimen that has been fully saturated using (3), the isotropic diffusivity with moisture could be extracted using nonlinear regression techniques

(1)

where is the moisture concentration. Note that hygroscopic swelling is a form of body force similar to thermal expansion. The hygroscopic swelling strain and stress can therefore be modeled using the thermal–structural functions of commercial finite-element software [3], [4]. C. Moisture Absorption and Diffusion Characterization The profile of moisture concentration within the underfill is needed for computation of hygroscopic swelling and can be obtained through moisture diffusion modeling [3], [4]. Two properties are required for such modeling: saturated moisture concentration Csat and moisture diffusivity . The saturated concentrations (Csat) of the underfill materials and the organic laminate (without solder mask) were charac-

(2)

(3) where

, , are the dimensions of the specimen and , LM , . Thin films of solder mask of 50- m thickness were specially prepared for the characterization. In view of the very thin dimensions and small volumes of the specimens, the characterization was performed on a specialized instrument that is equipped with sophisticated microbalance and temperature–humidity capabilities [13]. III. RESULTS A. T–H Aging The test results are tabulated in Table II. Note that: 1) the autoclave performance of the four underfill materials is similar to

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Fig. 5. Computation of CME for underfill UF-A at 85 C. (a) TGA measurement. (b) TMA measurement. (c) Combined TGA and TMA measurement.

Fig. 3. Autoclave failure signatures of flip chip PBGA assembled with different underfill materials.

Fig. 6. Linearity of concentration and swelling for underfill UF-A over a range of temperature.

and UF-D. Type II failure is typified by die-underfill delamination initiating at the die edge and propagating toward the center of the die, as observed for underfills UF-A and UF-C. B. Hygroscopic Swelling Characteristics of Underfill Materials

Fig. 4.

The two distinct failure modes in autoclave aging test.

that seen in earlier experiments (Table I), therefore indicating repeatable and consistent test data, and 2) the life of the packages at 85 C/85% RH aging is more than 30 times longer than the life at 121 C/100% RH aging. The scanning acoustic images of the packages assembled with the four underfill materials, taken before and after the autoclave test (at durations stated in Table II), are shown in Fig. 3. Two characteristic failure signatures were observed (Fig. 4). Type I failure is typified by random and localized delaminations along the die-underfill interface that subsequently coalesce into a large interior delamination, as observed for underfills UF-B

The TGA and TMA measurements of underfill UF-A at 85 C are presented in Fig. 5(a) and (b), respectively. The use of these two measurements to compute the CME is presented in Fig. 5(c). Note the near perfect linearity between the hygroscopic swelling and the moisture content. Fig. 6 shows that this linearity holds true over a wide range of temperatures. The same is also true for the other three underfill materials. Although the CME data at 170 C and 220 C are not needed in this analysis, they are nevertheless included to illustrate the temperature sensitivity of CME. Note that a thicker specimen has been used to characterize CME and diffusivity at these two temperatures in order to minimize error caused by the out-gassing of solvent as well as the depletion of moisture due to the long duration for temperature ramp-up [13]. The CME–temperature characteristics of the four underfill materials and the organic laminate are summarized in Table III. Observe that: 1) there is a wide variation among the CME of different underfill materials, ranging from 0.28 for UF-D to 0.45


CMES

OF

TABLE III UNDERFILL MATERIALS AND ORGANIC LAMINATE FUNCTION OF TEMPERATURE

AS A

TABLE IV MOISTURE ABSORPTION AND DIFFUSION CHARACTERISTICS UNDERFILL MATERIALS AND SUBSTRATE

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OF THE

Fig. 7. Extraction of moisture diffusivity using regression fitting for UF-A at 85 C. Fig. 9. UF-C.

A typical 1-D moisture diffusion modeling at 121 C/100% RH for

TABLE V MINIMUM HOURS REQUIRED FOR MOISTURE SATURATION OF PACKAGE

Fig. 8. Moisture diffusivity of the four underfill materials exhibit Arrenhius behavior.

for UF-B, and 2) the CME of underfill is weakly dependent on temperature, especially true for UF-A and UF-B. C. Moisture Absorption and Diffusion Characteristics of Underfill Materials A typical regression fitting to extract the moisture diffusivity from the TGA measurement is illustrated in Fig. 7 for underfill UF-A at 85 C. Similar extraction procedures have been applied for all the four underfill materials at four temperatures—85 C, 120 C, 170 C, and 220 C. Moisture diffusivity for the four underfill materials has been observed to obey the Arrhenius relation (Fig. 8). The absorption and diffusion characteristics of the underfill materials as well as that of the organic substrate used in the aging experiment are tabulated in Table IV. Note that while the Csat of the underfill materials at 121 C/100% RH is 15% higher than that at 85 C/85% RH—due to the linearity assumption—the diffusivity of the underfill materials at 121 C/100% RH is 2.0–2.5 times higher than that at 85 C/85% RH. The question therefore arises as to whether there is significantly more moisture absorption in the underfill in the autoclave experiment, due to the much increased diffusivity.

IV. ANALYSIS AND DISCUSSIONS A. Why Is 121 C/100% RH Aging Much More Severe Than 85 C/85% RH Aging? It is intuitive to suspect that hygrothermal strain is the reason behind the observed large disparity between the package reliabilities at the two aging conditions. To validate this conjecture, moisture content in the flip chip assembly under the two aging conditions was modeled, followed by computation of the hygrothermal strain in the underfill. Moisture Content in the Underfill: In the absence of a large copper plane, moisture diffuses predominantly through the organic substrate into the underfill, a one-dimensional (1-D) diffusion. A typical FE model for the underfill UF-C is illustrated in Fig. 9. The minimum number of hours to saturation for the four underfill materials at 85 C/85% RH and 121 C/100% RH have been tabulated in Table V, together with the time to failure. It turns out that with the exception of underfill UF-A at 121 C/100% RH which experienced 80% moisture saturation at failure, all the underfill materials experienced more than 95% moisture saturation at 121 C/100% RH and 100% moisture saturation at 85 C/85% RH prior to failure. In another words,

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TABLE VI THERMAL AND CURE PROPERTIES OF THE FOUR UNDERFILL MATERIALS

TABLE VII HYGROTHERMAL STRAIN IN THE UNDERFILLS

Fig. 10.

when failure occurs, the amount of moisture and the accompanying hygroscopic strain in the underfill at 121 C/100% RH are merely 15% higher than that at 85 C/85% RH. This appears to indicate that hygroscopic strain is independent of the moisture diffusivity of the underfill materials. However, could this extra 15% of hygroscopic strain have so severely reduced the life of the flip chip package? Hygrothermal Strain in the Underfill: The underfill in the flip chip assembly experiences three sources of strain during the T–H aging: shrinkage strain from curing of underfill, which is compressive in nature; thermal strain from temperature excursions between the assembly process and the T–H aging, which is usually compressive in nature; and hygroscopic swelling strain from moisture absorption, which is tensile in nature. A simplified 1-D free-expansion analysis was performed by computing the hygroscopic and thermal strain in the underfill materials during the T–H aging using (1) and (4), respectively

(4) where , are the CTE of underfill material at temperature and below and above its glass transition temperature is the cure temperature of the underfill at assembly. The properties of the four underfill materials are tabulated in Table VI and the computed hygrothermal strain in Table VII. Lack of appropriate information for the effective cure shrinkage—the term “effective” being used because only a fraction of the total volumetric shrinkage occurs in the “solid” state and results in “engineering strain”—does not allow the computation of cure shrinkage strain to be performed. Tensile hygrothermal strain occurs during the autoclave test, and the magnitudes of the hygroscopic swelling strain correlates well with the autoclave performance of the respective underfills, showing that large strains result in a low life. There were significantly lower magnitudes of hygrothermal strains under

Type I failure mechanism.

85 C/85% RH aging, with hygrothermal strains in UF-D becoming compressive in nature. However, the difference in the magnitudes of the hygrothermal strains of the two aging conditions does not reflect the large disparity in aging life. Moreover, it is more difficult to explain why UF-D, which has a hygrothermal strain at 121 C/100% RH that is four to six times lower than that of the other underfills at 85 C/85% RH, has life that is 25 times shorter. Stress Corrosion Cracking: The fact that most of the failures occurred after full moisture saturation was reached in the underfill suggests the existence of a dominant time-dependent failure mechanism such as stress corrosion cracking. Stress corrosion cracking, a form of environment-assisted cracking, is a subcritical flaw growth mechanism involving a cooperative interaction between a static stress and the environment [14]. Simply put, it is the propagation of a flaw under an applied energy release rate , made possible by that is below the critical value, i.e., the presence of corrosive agents in the environment. The strong influence of moisture on the rate of subcritical interface debonding has been convincingly demonstrated [15]. At a constant temperature of 28.5 C and under the same applied strain energy release rate, the subcritical interface debonding ) has been found to increase by three orders of magrate ( nitude when the humidity was raised from 4% RH to 98% RH. The effect of temperature on subcritical crack growth has also been demonstrated in bulk glass and has been found to obey the Arrhenius relation [16]. In another words, the rate of subcritical debonding in a package at 121 C/100% RH may be many orders of magnitude higher than that at 85 C/85% RH, for the same level of hygrothermal strain. It is therefore likely that at even at a lower magnitude of hygrothermal strain, the rate of subcritical debonding at 121 C/100% RH may still be many times higher than that at 85 C/85% RH, as is the case with UF-D (Table VII). B. Autoclave Aging Failure Signatures The mechanism for Type I failure is illustrated in Fig. 10. The far-field hygroscopic swelling of the underfill induces transverse tension in the solder joint which in turns induces shear and tension in the underfill around the solder. Over time, subcritical debonding occurs between the underfill and the die around the solder joint. Neighboring debonded regions then coalesce into a large delamination. The random nature of the initial debonding is attributed to the simple fact that every solder joint in the package (except those near the boundary) experiences a similar magnitude of transverse tension, as evident in the FE modeling (Fig. 11). The loss of adhesion between the underfill and


Fig. 11.

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FE model showing near identical transverse stress in the solder joints.

Fig. 12. Hygroscopic swelling mismatch induces interior debonding near center of chip.

!

!

Fig. 14. Illustration of a 2-D hygrothermal stress modeling for UF-A. (a) T (121 C). (b) Hygroscopic stress, Dry 100 Thermal stress, T(cure) RH. (c) Integrated hygrothermal stress. TABLE VIII CORRELATION OF SY/SXY RATIO WITH FAILURE MODE Fig. 13.

Type II failure mechanism.

die raises the level of transverse tension in the solder joints. Also, the in-plane hygroscopic swelling mismatch among the organic laminates, the underfill, and the inorganic chip further contributes to tensile loading on the solder joints. The mismatch induces bending moments in the package which are resisted by tension in the solder joints. Hence, the eventual failure of solder joints promotes bowing of the package. The differential stiffness between the chip and the laminate results in tensile stress across the interface and leads to extensive coalescence of delamination near the center of the chip as illustrated in Fig. 12. Type II failure exhibits extensive delamination initiating from chip corners and edges. The solder joint failure also exhibits the distance from neutral point (DNP) phenomenon which is usually associated with shear loading. In-plane hygroscopic swelling of the organic laminate and underfill is believed to be the main driver behind this type of failure (Fig. 13). What factors influence the occurrence of the two failure modes? To answer this question, a two-dimensional (2-D) hygrothermal stress analysis was performed for the four underfill materials. A typical simulation for underfill UF-A is illustrated in Fig. 14. The thermal stress in the package due to the temperature excursion from cure temperature to the autoclave test temperature was computed [Fig. 14(a)], followed by computation of the hygro-stress in the package due to autoclave aging [Fig. 14(b)]. These two computations were combined to give the hygro-thermal stress in the package during the autoclave test (Fig. 14(c)). With the postulated failure mechanisms for the Type I and Type II failures as a basis, the transverse stress in the bump as well as the shear stress at the die edge were extracted from the FE model for the four underfill materials and tabulated in Table VIII. The shape and size of the finite elements used in all the models were kept constant so that the ratio of the transverse stress at the bump (Sy) to the shear stress at the die edge (Sxy) is relatively independent of the finite-element shape and size.

The analysis seems to suggest a correlation between the Sy–Sxy ratio and the observed failure signatures, with a high Sy–Sxy ratio promoting the Type I failure mode. In view of the significant effect of solder flux residue on the adhesion strength of the underfill–die interface, it is very likely that the occurrence of Type I failure, initiating from local underfill–die debonding, would also depend on factors such as the quantity and characteristics of the flux residue as well as the nature of the solder flux to underfill interaction. C. Parametric Analysis A simple parametric analysis has been carried out to identify the desired underfill material and process characteristics. At 121 C, the four underfill materials used in this work have a maximum differential CME of 0.16 cm g (between UF-B of 47 ppm C (between and UF-D) and differential CTE UF-A and UF-C). This translates to an average differential hygroscopic strain of 1.8 10 and differential thermal strain of 0.7 10 for the four underfill materials. Reducing the CME of the underfill material is therefore a feasible and more effective way of enhancing the autoclave reliability of a flip chip assembly. The minor positive impact gained from increasing the of the underfill material needs to be balanced against CTE the heightened risk of thermal cycling failure. However, a simple and yet effective means of improving the autoclave performance of an underfill material is to raise the curing temperature of the underfill in the assembly process. Simply raising the curing temperature of the four underfill

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materials by 15 C would introduce an additional compressive thermal strain of 1.5 10 . This is accompanied by the extra benefits of: 1) enhanced thermal cycling fatigue life of the solder joint due to the additional compressive strain and 2) enhanced moisture resistance of the flip chip package due to a higher strain-free temperature of the package. V. CONCLUSION Under constant T–H aging such as the standard 85 C/85% RH, autoclave or highly accelerated stress test (HAST) tests, hygrothermal strains in the underfill would not result in instantaneous failure of the package. Rather, catastrophic failure of the package would occur after a period due to subcritical crack growth. Stress corrosion subcritical debonding of the underfill–die interface, at a rate that depends exponentially upon the applied energy release rate and temperature, is believed to be responsible for the observed exponential difference in the life at 85 C/85% RH and 121 C/100% RH aging. Two characteristic failure signatures have been observed in the autoclave failure of the flip chip–underfill packages: Type I, typified by random and localized delamination along the dieunderfill interface that could coalesce into a large interior delamination, and Type II, typified by die-underfill delamination initiating at the die edge and propagating toward the center of the die. The ratio of the transverse stress at the bump (Sy) to the shear stress at the die edge (Sxy) seems to correlate well with the two failure signatures of the packages, with Type I failure promoted by a high Sy–Sxy ratio. The autoclave performance of the flip chip–underfill could be enhanced more effectively by lowering the CME of the underfill material and less effectively by increasing the CTE of the underfill material, which carries the penalty of reduced thermal cycling resistance. Raising the curing temperature of the underfill is the simplest and the most effective means of improving the autoclave performance of a flip chip package. This is accompanied with the added benefits of enhanced thermal cycling and moisture resistance of the package. ACKNOWLEDGMENT The authors would like to thank S. W. Koh for the finite-element analysis performed and S. K. W. Seah for the excellent proofreading. REFERENCES [1] Y. Diamant, G. Marom, and L. J. Broutman, J. Appl. Polym. Sci., vol. 26, pp. 3015–3015, 1981. [2] M. R. VanLandingham, R. F. Eduljee, and J. W. Gillespie Jr, “Moisture diffusion in epoxy systems,” J. Appl. Polym. Sci., vol. 71, pp. 787–798, 1999. [3] E. H. Wong, K. C. Chan, T. Y. Tee, and R. Rajoo, “Comprehensive treatment of moisture induced failure in IC packaging,” in Proc. 3rd IEMT, 1999, pp. 176–181. [4] E. H. Wong, K. C. Chan, R. Rajoo, and T. B. Lim, “The mechanics and impact of hygroscopic swelling of polymeric packaging materials in electronic packaging,” in Proc. 50th Electronic Component Technology Conf., 2000, pp. 576–580. [5] J. F. M. Caers, R. Oesterholt, R. J. L. Bressers, T. J. Mouthaan, and J. F. Verweij, “Reliability of flip chip on board—First order model for the effect on contact integrity of moisture penetration in the underfill,” in Proc. 48th Electronic Component Technology Conf., 1998, pp. 867–871. [6] Y. W. Huang, K. H. Teo, K. L. Chua, M. W. R. Yang, and W. Ferng, “The effects of underfill on the pressure cooker test performance of flip chip on board assembly,” in Proc. InterPack, 1999, pp. 1121–1127.

[7] E. H. Wong, S. W. Koh, R. Rajoo, and T. B. Lim, “Underfill swelling and temperature–humidity performance of flip chip PBGA package,” in Proc. EPTC, 2000, pp. 258–262. [8] E. G. Wolff, “Polymer matrix composites: Moisture effects and dimensional stability,” in International Encyclopaedia of Composites. New York: VCH, 1991, vol. 4, pp. 279–323. [9] Test Method for the Measurement of Moisture Diffusivity and Water Solubility in Organic Materials Used in Integrated Circuits, JEDEC Std. JESD22-A120, 2001. [10] C. H. Shen and G. S. Springer, “Moisture absorption and desorption of composite materials,” J. Compos. Mater., vol. 10, no. 2, pp. 2–20, 1976. [11] A. C. Loos and G. S. Springer, “Moisture absorption of graphite–epoxy composite immersed in liquids and in humid air,” J. Compos. Mater., vol. 13, no. 131, pp. 131–147, 1979. [12] E. H. Wong, S. W. Koh, K. H. Lee, and R. Rajoo, “Advanced moisture diffusion modeling and characterization for electronic packaging,” in Proc. 52nd Electronic Component Technology Conf., 2002, pp. 1297–1303. [13] E. H. Wong and R. Rajoo, “Moisture absorption and diffusion characterization of packaging materials- advanced treatment,” Microelectron. Rel., no. 43, pp. 2087–2096, 2003. [14] R. W. Hertzberg, Deformation and Fracture Mechanics of Engineering Materials, 4th ed. New York: Wiley, 1996, ch. 11. [15] S. Y. Kook and R. H. Dauskardt, “Moisture-assisted subcritical debonding of a polymer/metal interface,” J. Appl. Phys., vol. 91, no. 3, pp. 1293–1303, 2002. [16] A. Bhatnagar, M. J. Hoffman, and R. H. Dauskardt, J. Amer. Ceram. Soc., vol. 83, no. 585, pp. 2000–2000.

E. H. Wong received the M. Eng. degree in solid mechanics from the University of Manchester Institute of Science and Technology (UMIST) in 1989. He spent a decade in the defense industry working in product and process development. He joined the Institute of Microelectronics (IME), Singapore, in 1996 and is currently a Senior Member of Technical Staff with the Microsystems Modules and Components Lab. He is also doing postgraduate study at the University of Sydney, Sydney, NSW, Australia. He has been active in modeling and material characterization.

Ranjan Rajoo received the Diploma degree in mechanical engineering from Singapore Polytechnic in 1999. He is currently Research Officer with the Microsystems Modules and Components Lab, Institute of Microelectronics, Singapore. He has been active in moisture and mechanical characterization of electronics packaging materials.

Thiam Beng Lim (M’92) received the Ph.D. degree from the University of Aston in 1994. He was with Texas Instruments in packaging development for memory devices. He is currently with the Institute of Microelectronics (IME), Singapore. He is an inventor/coinventor on 21 U.S. patents and authored/coauthored some 50 publications in electronic packaging journals and conferences. Dr. Lim is an Active Member of the IEEE Components, Packaging, and Manufacturing Technology (CPMT) Society. He is a Founding Organizing Committee Member of the annual IEEE/CPMT Electronics Packaging Technology Conference in Singapore. He was awarded the IEEE CPMT Society Outstanding Sustained Technical Contributions Award in 2003.

Yiu-Wing Mai received the Ph.D. and D.Sc. degrees from Hong Kong University in 1972 and 1999, respectively, and the D.Eng. degree from the University of Sydney, Sydney, NSW, Australia, in 1999. He is University Professor and ARC Federation Fellow at the University of Sydney. He is also Director of the University’s Center for Advanced Materials Technology (CAMT). He is a highly cited author in materials science for two decades (1981–1999), listed at http://www.isihighlycited.com/. His research interests are fracture and fatigue mechanics, tribology, and processing–structure–property relationships of a range of engineering materials including polymers, composites, and thin-film and nanostructured materials. Prof. Mai is a Fellow of the Australian Academy of Science and Australian Academy of Technological Sciences and Engineering, as well as the Hong Kong Academy of Engineering Sciences and the American Society of Mechanical Engineers.