An Energy-based Method to Predict Delamination in Electronic Packaging

H.B. FAN, Paular W.K. CHUNG, Matthew M.F. YUEN, and Philip C.H. CHAN* Department of Mechanical Engineering,

and Department of Electrical and Electronic Engineering", Hong Kong University of Science and Technology Clear Water Bay, Kowloon, Hong Kong SAR, China

Email: mehaibo@ust.hk, mepaular@ust.hk, meymf@ust.hk, eepchan@ust.hk Phone: (852) 2358 8814; FAX: (852) 2358 1543

Abstract The propensity and significance of interfacial

delamination as a crucial failure mechanism in electronic packaging have been well documented in many papers. Many of the failure criteria were used to solve 2-dimensional problem with a pre-crack. However, in real electronic packages, the size and location of the cracks orland delamination cannot be predicted. It is not easy to use the traditional fracture criteria to deal with more complicated 3-D delamination problems. The potential delamination interface of copper leadframe/Epoxy Molding Compound (EMC) was selected in the study. The stresses of the interface were evaluated by the Button Shear Test.

A series of Button Shear Tests was conducted to evaluate the adhesion properties of Epoxy Molding Compounds (EMCs) on copper substrate. In each of the tests, the critical load acting on the EMC of the button shear sample was measured at different shear angles and a finite element model was used to evaluate the stresses at the interface between the mold compound and the copper substrate. In this paper, an energy-based method is proposed by deriving the energy to initiate each of the tensile and shear modes of failure across the interfaces of the button shear test samples for the chosen EMCfleadfiame material system. Component stresses were extracted from the numerical simulation in order to compute

the distortional strain energy density, , and the

hydrostatic strain energy density, , relating respectively to

the shear and tensile mode. and were calculated from the Young's modulus of EMC and the average stresses within a selected region of the finite element model where it exhibits high stress values.

Introduction Interfaces in electronics packages are susceptible to

delamination due to demanding packaging requirements, stringent qualification and working environment. Delamination, especially those occurring between copper (Cu) and Epoxy Molding Compound (EMC) in plastic packages, often leads to popcoming during manufacturing processes.

The propensity and significance of interfacial delamination as a crucial failure mechanism in electronic packaging is well documented. Generally, two different failure criteria have been commonly employed to characterize the interfacial strengths of composites and IC packages. One of them is based on the fracture mechanics approach, where the interface crack propagates when the interface stress

intensity factor or energy release rate exceeds the critical value [l]. This approach is useful for analyzing the reliability and life of a product that is assumed to contain some inherent cracks. However, the amount, location and size of delamination or crack cannot be easily determined in Ieal packages, it thus poise a problem for predicting failure at interfaces in IC packages.

The other criterion is the strength-based approach, where some predefined critical loads or average stresses are utilized to assess the interface strength [2-51. In this approach, the maximum load that materials or structure can sustain before interface failure is measured by applying tensile or shear loads to the test specimens. It is relative easy to perfcirm experiments based on the strength-based approach and the failure criterion developed on the basis of the test will be very useful for design of products and the prediction of product reliability.

Among of the strength-based approach, strain energy density based criteria assume that failure occurs when the strain energy density at any location in the interface exceeds a limiting value. This limiting value can be defined as the area under a tensile or shear stress-strain curve. The criterion was introduced by Sih and Macdonald [6] and later applied to an extensive class of problems by Sih and his co-workers [7- 81. Based on an ideal elastic-plastic and bilinear plastic model, Hart-Smith [lo] proposed a simple formula to express the failure load of an adhesive bonded double lap joint in terms of the strain energy density.

In the present study, the Button Shear Test is employed to evaluate the interfacial adhesion between EMC and copper. Finite element modeling of the button shear test samples was conducted to obtain interfacial stresses. Average strain energy density was then calculated. The failure criterion was verified experimentally with five different EMC materials.

Button Shear Test The button shear test, which is commonly adopted for

EMC testing, was carried out to evaluate the bi-material properties between EMC and copper leadframe. The leadframe material investigated has a 99.8% copper contmt. Five commercial EMCs were selected for the study.

Copper strips with the dimension of 13" wide itnd 70" long were prepared for the study (a schematic diagram is shown in Figure 1. The copper leadframe strip was degreased, cleaned using oxide cleaning agents and dry cleaned under UV-ozone for 5 minutes. The button-shape EMC was then molded on the cleaned copper strip and post

0-7803-7430.41021$17.00 02002 IEEE 834 2002 Electronic Components and Technology Conference

cured subsequently. The curing process follows the specification of manufacturers.

EMC 1

EMC 2

EMC 3

EMC 4

EMC 5

c u Figure 1: The dimension (in mm) of the button shear sample.

Shear tests were performed using the Dage 4000 shear tester with a DSlOO load cell. The button shear test specimen was fixed on the mounting jig attached onto the shear tester platform. A special testing jig was designed and made for clamping the sample and to cater for testing with preset angular loadings of -lo", On, 10" and 20". The schematic showing the angular jig set up is shown in Figure 2.

A loading rate of 85pm/s was applied to the button surface with a controlled loading height, as shown in Figure 2. The fracture surface of each sample was inspected. The breaking load of the specimen with interfacial fracture was recorded. Five samples were tested for each loading height setting. All experiments were performed at room temperature.

Young's Modulus Poisson's Ratio (G Pa>

18.2 0.3

14.2 0.3

0.3 23.8

13.3 0.3

17.2 0.3

110 0.36

H

2O%H Angle J 7

A three-dimensional finite element model was built to calculate the stresses at the EMC/Cu interface. Eight-node solid element was used. The global finite element meshing is given in Figure 3. The mesh was refined at the edges and interface to cater for the steep stress gradients expected as shown in Figure 4. The elements varied in sizes from 0.06 X 0.12 X 0.1" near the interface edge to 0.8 X 0.05 X 0.1" near the interface center. Boundary conditions and loading were applied to emulate the test conditions in the button shear test. The displacements of nodes at the bottom surface of the copper substrate were constrained.

Figure 2: The schematic drawing of the angular shear test.

Finite Element Model In order to study the interfacial delamination of EMC and

Cu under mechanical loading, it is necessary to perform a finite element analysis to extract the stresses. Due to the complexities associated with a bi-material wedge of finite dimensions, numerical finite element stress analysis using the ANSYS code was employed to all test specimens. The model in this paper consists of EMC and copper. The geometry is shown schematically in Figure 1. The materials are assumed to be linear elastic, homogeneous and isotropic. The Young's modulus and Poisson's ratios of the materials are listed in Table 1.

Figure 3: The finite element model of Button Shear Test

Approach In order to predict the interfacial delamination initiation

and growth, it needs to consider the stresses and strains across the interface for all test conditions. Several failure criteria for interfacial failure were based on the strength- based fracture mechanics approach. Sih [8] developed the Strain Energy Density (SED) failure criterion that examines the local strain energy density in the region of interest. This criterion has been successfully used in two- and three- dimensional crack problems; dynamic crack problems; failure initiating from notches; ductile fracture involving the prediction of crack initiation; slow stable crack growth and final separation, etc. For linear elastic material behavior, the

835 2002 Electronic Components and Technology Conference

total strain energy density and distortional strain energy density at various locations can be obtained using equation (1) and equation (2) respectively. The hydrostatic strain energy density can be expressed as equation (3).

EMC 2

(3)

Where U , and are the total, distortional and hydrostatic strain energy per unit volume respectively: CY and z are the stress components in the x, y and z directions; E is Young’s Modulus; v is Poisson’s ratio.

742.1 639.4 545.18 474.26

-litbpirriru

(b) Figure 4: The (a) top and (b) front views of the meshed button and substrate of the finite element model.

EMC 3

In this approach, the values of strain energy density of the interfacial elements are averaged across the EMC/Cu interface. The strain energy values of each element are normalized by the volume of the, selected interface elements. The average of strain energy density can be written as:

479.79 416.85 338.19 304.74

Where is the average strain energy density for the interfacial elements, U i is the strain energy density of each element, and Vi is the volume of each element.

According to the above formula, the average distortional and hydrostatic strain energy densities are c d and t i h

respectively, and total strain energy density could be calculated by equation (3).

Results and Discussion

Experimental Results The specimens were divided into 5 groups with 5

different EMCs and one copper material. In each group, 4 different shear angular loadings will be applied in the shear test for one material system. A typical load-displacement graph was obtained from the Angular Button Shear Test as shown in Figure 5. It can be seen that the loadirig- displacement relationship is fairly linear. The average failure loads obtained with different angles for five various materials are listed in Table 2. As expected, the critical shear forces become smaller with angle changing from -10’ to 20g for all EMC materials. All experimental results show a clean substrate surface after debonding indicating interfacial adhesion failure in the Angular Button Shear Test.

EMC4 358.75 302.60 222.12

300

250 5. 200

p 150 3

100

50

0

c

199.84

I

EMC 5 650.32 449.15 379.83

0 0.05 0.1 0.15 0.2 Displacement (mm)

328.68

Figure 5: A typical load-displacement curve of the Button Shear Test.

Table 2: Critical shear forces (N) for different materials. hear angle ‘i Mterial -102

591.25 413.23 378.43 333.31

I I I I I

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Finite Element Results The failure load for each case was applied to the finite

element stress analysis. There are complex stress states and stress concentrations at the free edge of the bi-material interface near the loaded edge of EMC, and the stress distribution of normal and shear stress at the interface from the free edge, that is the stress singular point, is shown in Figure 6 using d as a distance from the free edge. Both the interfacial normal and shear stresses are of the greatest magnitude at the free edge, so it is confirmed that debonding initiation will occur at the edge.

5 0 -

0 -

8 EMC 2 A EMC 3 - EMC 4 X EMC 5

P" 150

; 100 - - - - L A

D 0.0005 =O.&l 0!!015' E 0 2 O.d?5 0.003

8 EMC 2 A EMC 3 - EMC 4 X EMC 5

P" 150

8 EMC 2 A EMC 3 - EMC 4

60 X EMC 5

50

0 4 I 0 0.0005 0.001 0.0015 0.002 0.0025 0.003

d fm) (b)

Figure 6: Distribution of the normal (a) and shear (b) stresses along the interface for different EMC materials

Discussion The interfacial delamination between the EMC and Cu is

mainly determined by the normal and shear stress distribution near the loaded edge of EMC. Due to the singular nature of the stresses at the free edge, the average stress approach was used for deducing the interfacial debonding criterion taking into account both the normal and shear stresses. Sun and Zhou [ l l ] and Kim and Soni [12] have proposed criteria to predict interfacial delamination at free edges of composite laminates based on averaging of interfacial stresses. To calculate averaging interfacial stresses, nodal stresses at the interface within the distance ro = 0.58mm around the free edge. The average interfacial normal and shear stresses data for five EMC materials are included the criterion, as given in Figure 7. It is obvious that the failure mode will change with the shear angle.

1 ~ E M C 11 1.2

1

0.8

13" 0.6 B 0.4

0.2

0

' mEMC 2

AEMC3 -EMC4 XEMC5

Increasing shear angle __

0 0.2 0.4 0.6 0.8 1 1.2

ZIZC

. Figure 7: Mixed mode of failure criterion

Due to stress concentration at the free edge, the interfacial strain energy density is always highest at this point. The strain energy density along the free edge is illustrated in Figure 8. Based on the results, it is assumed that delamination occurs at the leading edge of the interface.

+EMC 1

8 EMC 2 AEMC3 -EMC4 X EMC 5

$ 0.6 ai

.; 15 m o 0.2 0.4 ~

0 0.0005 0.001 0.0015 0.002 0.0025 0.003

don)

Figure 8: Distribution of Strain Energy Density along the interface for five EMC materials.

Across the interface, the Strain Energy Density criterion states that delamination for the particular bimaterial system may initiate at the interface if the average value of SED across the interface exceeds a certain critical value. In order to average the SED around the edge, two adjacent layers of interfacial elements were selected across the two sides of the interface within the distance ro = 0.58" around the free edge. Under different loading condition, the critical average distortional strain energy density c d and hy"Tostatic strain energy density e h were calculated for the different EMC/copper material systems respectively. The results for five EMC materials are shown in Figure 9. Clearly, it can be seen that c d decreases and e h increases with the shear angle changing from -10" to 20" and the magnitude Of c d is higher than that of c h . It can be concluded that c d has more contribution to interface debonding when shear angle is at - 10" and on the contrary c h has more contribution to the interface debonding as shear angle becomes larger. Because the critical value of e h at the interfacial edge is much lower

837 2002 Electronic Components and Technology Conference

than that of I?d , the energy for interface debonding becomes rather small as the shear angle becomes larger. Likewise the failure load will become smaller with a larger angle.

35000 - Increasing shear angle - B 30000 -

25000 - ’\\ -‘-\,

*EMC 1 B EMC 2 AEMC 3 - EMC 4 X EMC 5

0 20000 40000 60000 80000 100000 120000 Ud

Figure 9: Averaging strain energy density Ud Vs Uh for 5 EMC materials

20000

15000 -

10000 -

5000 -

0,

c 3

1.2

1

0.8 U c 2 0.6 3

0.4

0.2

0

\ - X 4

A x* -A A X * -

-A X 7

I E M C ~

EMC 2 AEMC 3 - EMC 4

X EMC 5

Increasing shear angle

0 0.2 0.4 0.6 0.8 1 1.2

Ud/Ud,

Figure 10: Mixed Strain Energy Density failure mode

Based on the above results, E d and E h are used here as two failure parameters to predict interfacial delamination. Strain energy density criteria is given as follow:

5 ( 5 ) +-=1

Where and U,,, are the critical distortional strain energy density under pure shear and the hydrostatic strain energy density under hydrostatic pressure, which are properties of the particular interface.

The failure criterion using the Strain Energy Density method can simplify the failure prediction of package design by using the Button Shear Test data for material testing to determine the critical Strain Energy Density. With the known value of Young’s modulus of EMC and average interfacial stresses, and U,, can be calculated in the real package numerical model. If the result falls outside the safe region defined by the failure criterion as in Figure 10, then interfacial failure may occur in the package.

Summary I

By using the Angular Button Shear Test, adhesion properties can be investigated easily. A new methodology was established to interpret data from the Button Shear Test

using the numerical simulation model to generate the Strain Energy Density interfacial failure criterion. By applying the finite element analysis to a real package model, imd determining the maximum distortional strain energy density and hydrostatic strain energy density at the interface, potential failure across material interface can be predicted.

Acknowledgments The project was supported by the Innovation and

Technology Commission project (AF/222/98) - “A program for delamination control in electronic packages“ and the Research Grant Council project (HKUST6104/97E) - “Efiect of thermal cycling on interfacial cracking in electranic devices“. The authors would like to thank Nitto Denko for supplying the EMC materials.

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838 2002 Electronic Components and Technology Conference