- Email: [email protected]
ILB interconnection reliability
Microelectronics Reliability 43 (2003) 935–943 www.elsevier.com/locate/microrel
Numerical study on the bonding tool position, tip profile and planarity angle influences on TAB/ILB interconnection reliability D.S. Liu
a,*
, Y.C. Chao a, C.H. Lin a, G.S. Shen b, H.S. Liu
b
a
b
Department of Mechanical Engineering, National Chung Cheng University, 160, San-Hsing, Ming-Hsiung, Chia-Yi 621, Taiwan, ROC ChipMOS Technologies Inc., No. 5, NAN-KO Road, 7, Science-Bases Industrial Park, Tainan 744, Taiwan, ROC Received 23 September 2002; received in revised form 5 March 2003
Abstract Tape automated bonding (TAB) is a widely used interconnection technology for high-pincount and fine-pitch IC packaging. In this study, a three-dimensional computational model was developed for analyzing TAB inner lead bonding (ILB) process. This experimental study on the thermomechanical properties of copper leads was achieved using high precision micro-force tensile tests. A stress–stain relation between the copper lead and different temperature ranges was successfully implemented into the finite element model to study large plastic deformation in ILB formation. The resulting ILB lead profile and bump sinking values obtained from the simulations agreed well with the experimental observations from actual manufacturing data with the same bonding parameters. The tool position and lead length effects are analyzed to study the residual stress distribution after ILB. A 10-lead model was developed to study how the tool tip profile and planarity Ôangle affect the co-planarity between the bonding tool and the stage. The numerical results show that the permissible tool profile variance should not exceed 1.25 lm and the acceptable planarity angle is 0.005 to achieve the minimum bump deformation requirement. 2003 Elsevier Science Ltd. All rights reserved.
1. Introduction Tape automated bonding (TAB) technology is a momentous interconnection method in first level packaging for portable and handheld electrical production. TAB offers many advantages such as finer pitch and high I/O, lower profile, more precise geometry and gang bonding. This assembly technology corresponds well with the current trend in electronics [1]. The first assembly step in the TAB process involves bonding silicon chips to patterned metal-on-polymer tape, called inner lead bonding (ILB). The objective of ILB is to form a strong metallurgical bond between the gold bump on the
*
Corresponding author. Tel.: +886-5272-0411x33305; fax: +886-5272-0589. E-mail address: [email protected] (D.S. Liu).
die pad and the tin plated beam lead on the carrier tape (see Fig. 1). This process does not compromise the quality and electrical reliability of the final component. Thermocompression (T/C) gang bonding is the most commonly used TAB/ILB technology. This method can simultaneously bond all TAB beam leads to gold bumps on a chip using a bonding tool (thermode, see Fig. 2) with proper temperature, pressure and dwell time. This bonding method does increase throughput, but on chips with high-density I/O, planarity problems can lead to missing bonds, cracked chips or inconsistent pressure. Improper tool temperature, bonding force and tool position can cause fractures in the beam lead span near the bond pad (see Fig. 3). Currently, the packaging industry relies heavily on the use of tests to adjust the bonding parameters to meet the design requirements. This trialand-error process is time consuming and cost ineffective. Therefore, it is important to fully understand the
0026-2714/03/$ - see front matter 2003 Elsevier Science Ltd. All rights reserved. doi:10.1016/S0026-2714(03)00095-7
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D.S. Liu et al. / Microelectronics Reliability 43 (2003) 935–943
Fig. 1. Schematic of ILB.
Fig. 2. Thermocompression bonding tool.
Fig. 3. Inner lead fracture.
dynamic characteristics of the TAB/ILB to develop better design tools that can quickly obtain optimum manufacturing parameters for various types of TAB packaging. In the past, much research has been performed to examine the possible TAB packaging failure modes due to thermal mismatch or cyclic thermal loading [2–7]. Few research studies attempted to investigate how the bonding parameters affect bonding strength and reliability. In experimental work, Atsumi et al. [8] studied the acceptable bonding conditions and the tool tip configuration for both the 182-lead and 504-lead test dies. Optimal bonding ranges for tool temperature, bonding pressure, and bonding time were found by examining the pull test bond strength. Lai et al. [9] investigated how bonding temperature and bonding pressure affected the microstructure of ILB joints. Oshige and Nakanishi [10] applied a simple two-dimensional numerical model to simulate the contact between a lead and bump pair. Only part of the lead close to the bump was modeled. Numerical results such as the deformation shape and strain distribution for the lead and bump were presented. Jung et al. [11] used the finite volume method to calculate the temperature distribution for each component during ILB, and concluded that the temperature of the heat cartridge in the bonding tool should not be above 520 C to avoid thermal damage (delamination) in the adhesive layer. Takahashi et al. [12] employed twodimensional contact finite element (FE) analysis to study the size ratio effect between the pad thickness and lead height in the ILB process. They found that if the pad thickness was decreased, the stress exhibited a peak under the pad that might cause damage to the die. Up to now, no research has presented a numerical model to examine how the geometric parameters such as the tool compression location affect the stress distribution on the beam lead span. Most studies used only a two-dimensional simple FE model to perform the analysis. This approach cannot be used to discuss the tool planarity and flatness effects. Therefore, in this research the TAB/ILB formation phenomena during the T/C bonding process was simulated using a single-lead threedimensional computational model and explicit FE analysis. Sn-plated Cu lead thin-strip samples were used as test specimens to perform micro-force tensile tests. The elasto-plastic stress–strain relations for the beam lead at various temperature ranges were combined into nonlinear dynamic FE code LS-DYNA3D to perform TAB/ILB formation simulations. The numerical results were used to study lead deformation, residual stress at the lead span after bonding and gold bump-downward length. A 10-lead FE model was also developed based on a single-lead model to study the effect of the tool tip profile and the tool and die coplanar surface. Both FE models were verified using actual TAB/ILB tests. The numerical simulation model presented here can provide
D.S. Liu et al. / Microelectronics Reliability 43 (2003) 935–943
better understanding of the fundamental formation phenomena during the ILB process and serve as a useful tool for better adjusting the bonding tool parameters to achieve the reliability requirements.
937
450
R.T. 400
100 oC 350 o
150 C
2. Copper lead constitutive model To determine the appropriate elasto-plastic constitutive model to quantify the plastic flow characteristic effects on the Cu lead in the ILB process, thermomechanical micro-force tensile testing was developed. The test sample material investigated in this study was electrodeposited copper film used as the raw material for Snplated Cu beam lead. Experiments were performed on an MTS-Tytron micro-force tester equipped with a 125 N load cell. An environmental chamber was attached to the testing machine to conduct thermal testing. The experimental setup is shown in Fig. 4. Copper films were cut into miniature thin-strip specimens with 40 mm · 5 mm (length by width) dimensions with 18 lm uniform thickness. To obtain consistent and accurate test results, the grip force was suitably adjusted to prevent the film from slipping or breaking. Specimen alignment was performed carefully to prevent warpage. An additional multi-axis translation stage and CCD camera was therefore used to provide the required motion and focus vision to ensure accurate alignment. Fig. 5 shows the tensile stress–strain curves at various temperature ranges. All of these curves appear the same where strain hardening occurs through the initial 0.2–2% strain and is followed by continued flow at minor stress increments. Examination of each curve in Fig. 5 illustrates that the strength decreases as the imposed temperature is increased. The maximum stress at room
Stress(MPa)
300 250
200 oC
200
250 oC 150 100
Copper Film Tensile Test
50 0
0
1
2
3
4
5
6
7
Strain(%)
Fig. 5. Stress–strain curves for copper film at strain rate 1 min1 .
temperature (440 MPa) is dramatically reduced to 150 MPa with the temperature increased to 250 C. An elastoplastic constitutive model with exponentiallaw form was proposed by the authors to describe the behavior of gold bonding wire [13]. The model was modified to adequately describe the behavior of the copper beam lead without strain hardening effects. The constitutive relations are written as follows: r ¼ E0 ð1 eae Þ E0 ðT Þ ¼ 481:3061 1:1775T 0:00046946T 2
ð1Þ T unit: C ð2Þ
aðT Þ ¼ 125:677 0:1331T þ 0:000694T 2
T unit: C ð3Þ
where E0 , a are the material constants. The first term in Eq. (1) expresses the elastic and transition portion of the stress–strain curve. The second term in Eq. (1) denotes the stress–strain relationship in a fully plastic range. A comparison of the predictions from Eq. (1) is presented in Fig. 6. The results calculated from the constitutive model agree well with the experimental data. 3. Finite element modeling and validation
Fig. 4. Schematic of experimental setup for copper film tensile test.
The analysis model configuration shown in Fig. 7 is based on the actual TAB/ILB process. Because of the symmetrical nature of the problem in hand, only half of the ILB model was used for the simulation. The dimensions of the gold bump were 100 lm · 40 lm · 17
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D.S. Liu et al. / Microelectronics Reliability 43 (2003) 935–943 500 450
R.T.
400
o
100 C
Stress(MPa)
350 o
150 C
300 250
o
200 C 200
250 oC
150
Copper Film Tensile Test
100
Model Prediction
50 0
Experimental data
0
1
2
3
4
5
6
7
Strain(%)
Fig. 6. Comparison of model prediction with experimental data of stress–strain curve of copper film.
lm. The thickness of the copper lead was 18 lm. The length of the copper lead beam was 274 lm. The FE model for analyzing the single lead TAB/ILB process is shown in Fig. 8. It consists of five major parts: the gold bump, the polyimide, the copper lead, the adhesive and the bonding tool. The FE mesh was created with a combination of 3-D element types. The gold bump, polyimide, copper lead, and adhesive were modeled as solid elements and the tool with rigid shell elements. The number of total solid elements is 4752, and 360 for the shell elements. The bonding tool temperature and dwell time during the ILB process is a crucial factor in determining the temperature distribution for each part of the ILB simulation model. Jung et al. [11] used the finite volume method to calculate the temperature distribution at various TAB tape areas with varying tool temperatures
and dwell time. In our actual TAB/ILB test, the bonding tool temperature was set at 500 C. The dwell time was set at 1.0 s. According to Ref. [11], the gold bump and lead were heated to about 220 C. Fig. 9 shows the stress–strain curve for the copper lead obtained from above and the gold bump from Ref. [13]. The material properties of the other parts are listed in Table 1. The FE code LS-DYNA3D was used in this study to simulate TAB/ILB transient formation process. In LSDYNA3D, the transient analysis is performed using an explicit direct-time-integration procedure and thereby avoids the need for matrix evaluation, assembly and decomposition at each time step as required by many implicit time-integration algorithms. The updated Lagrangian (UL) FE formulation adopted in this code accounts for both the material and geometric nonlinearity. Using this formulation the FE mesh is continuously updated during the calculation, and all variables calculated from UL FE are referred to the last calculated configuration. Piecewise linear elasto-plastic material model was used to input the stress–stain curves for copper lead and gold bump. The other parts of the TAB were modeled by elastic material model. Loading curve then input as a mechanical loading to simulate bonding tool pressure applied to inner lead and gold bump. The bonding tool speed was 5 mm/s downward. Some special techniques for TAB/ILB process FE simulation, such as the contact interfaces, must be carefully defined between tool and lead and an element-tied criterion based on the maximum contact force between lead and bump, are proposed and implemented. Once the contact force between the lead and bump reached the bonding pressure limit for a single lead element on the bottom surface of the lead, the top elements on the bump are tied. To validate this FE model, we simulated the TAB/ ILB process with various bonding pressures ranging from 20 to 45 gf and compared the lead profiles and gold bump sinking values with actual ILB test data. Fig. 10 compares the ILB lead deformation profile with a
R40
274 18
33
12
17
434 100
130
Fig. 7. Configuration of single lead model (in lm).
40
120
75
70
D.S. Liu et al. / Microelectronics Reliability 43 (2003) 935–943
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Fig. 8. Single lead FE model: (a) oblique view and (b) side view.
250
Table 1 Material properties of single lead model
Stress(MPa)
200
Copper Lead Gold Bump
Adhesive Polyimide Tool
150
Elastic modulus (GPa)
PoissonÕs ratio
0.6367 3.15 Rigid
0.28 0.3
Yield stress (GPa)
Plastic hardening modulus (GPa)
0.085
0.97
100
50
0
0
1
2
3
4
5
6
7
8
9
Strain(%)
Fig. 9. Stress–strain curve of copper film and gold bump.
bonding force test results equal to 30 gf. Fig. 11 shows a schematic of the bump sink value definition and compares the bump sink values with the test results at various bonding pressures. Both agree very well with the test results.
Fig. 10. Comparison of the deformed profile of copper lead (upper: test results, lower: numerical results).
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D.S. Liu et al. / Microelectronics Reliability 43 (2003) 935–943 3.5 Experiment Simulation
Sinking Value(µm)
3
2.5
2 A
C
B
1.5
1
0.5
Fig. 13. Maximum Von-Mises stress.
A : L ead Thickness B : Bump Height C : Measurement Sinking Value = (A-B)-C
20
30
40
50
Bonding Force(gf)
Fig. 11. Comparison of numerical data with experimental data of gold bump sinking value.
4. Single lead analysis and results 4.1. Failure mode of beam lead and bonding force effects
Fig. 14. SEM micrograph of fracture of a beam lead.
240 220
Max. Von Mises Stress(MPa)
From the simulation the TAB/ILB process might be separated into four steps, as shown in Fig. 12. Step one, the tool makes initial contact with the end of the lead beam. Step two, the lead slides along the tool tip and makes contact with the bump-top surface. The sliding distance is about 50 lm. Step three, the tool compresses the lead top surface and ties the lead bottom surface with the bump. Step four, the tool rises up after completing the ILB process. Fig. 13 shows the effective (VonMises) stress contour distribution of the deformed beam lead after the third step. The maximum effective stress value occurred at the elements close to the tool tip. This observation is consistent with the failure mode from actual TAB/ILB tests, as shown in Fig. 14. Fig. 15 shows the maximum effective stresses vs. various bonding forces at the beam lead near the tool tip and in the gold bump. A larger bonding force increases
200 Inner Lead Gold Bump
180 160 140 120 100 80 60 40 20 20
25
30
35
40
45
Bonding Force(gf)
Fig. 15. Max effective Stress on lead and gold bump.
Fig. 12. ILB process.
the effective stress in the gold bump and copper lead. Excessive bonding force might be considered to be one of the driving forces causing damage to the silicon chip. Conversely, insufficient bonding force decreases the bump deformation and might produce a poor mechan-
D.S. Liu et al. / Microelectronics Reliability 43 (2003) 935–943
Effective Plastic Strain(%)
6
Lead Length 0.274mm Lead Length 0.324mm Lead Length 0.374mm
5
4
3
2 -20
0
20
40
60
Tool Position(um)
Fig. 16. Effective plastic strain on tool lip lead.
ical or electric bond. In actual TAB/ILB tests, the bonding tool force was set at 30 gf. In the following parameter study all results were calculated under bonding force equal to 30 gf. The bump deformation was the criteria for good bonding. According to the manufacturerÕs pull test data, a bump deformation range from 1 to 7 lm represents a good connection.
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gang bonding, especially as chip sizes become larger and the bump and lead geometry become smaller. In addition to the bonding tool planarity setup, the profile of the bonding tip surface at various elevated temperature is also critical to a successful gang ILB process. As the temperature increases up to 500 C, the profile of the tool tip surface exhibits a concave shape with approximately 1 lm of deformation at the periphery of the tip. A 10-lead FE model was developed to study the tool profile variation and planarity problems in the ILB process. The numerical results were compared with the experimental data. Fig. 17 shows the 10-lead FE model. The distance between the two leads is 50 lm. The material properties are the same as that for a single lead model. The tool surface was approximated as a concave circular arc that could be defines as long as the tool planarity distance (DH ) and chip length are given (see Fig. 18). To verify the accuracy of the computational model, simulations were carried out with DH ¼ 1 lm and chip length equal 16.85 mm according to the actual TAB/ ILB bonding experiments. The actual bonding samples were bonded using a Shibaura TAB bonder. The bump deformation for each sample was measured using a scanning electron microscope (SEM). Each experimental result was averaged from five testing specimens. Fig. 19 compares the numerical results with the gold bump sink
4.2. Bonding tool position / lead length effects The original horizontal distance between the tool tip and polyimide was 274 lm. This was chosen as the base line position (zero position). Four additional tool positions were examined; positions 1 to 3 involved moving the bonding tool horizontally with respect to the zero position 20, 40, and 60 lm toward to the edge of the polyimide. Position 4 involved moving the tool inversely 20 lm ()20 lm position) with respect to the original tool position. The single-lead FE model was performed to analyze these four cases to understand how the tool position affected the residual stress/strain. In addition, the length of the lead beam was increased to 324 and 374 lm to study the lead length effect. Fig. 16 illustrates the effective plastic strain on the lead near the tool tip. Increasing the beam lead length lead to a decrease in the effective plastic strain. Moreover, the slope of the curves in Fig. 16 decreased with increasing beam lead length, indicating that a shorter beam lead is more sensitive to the tool position.
Fig. 17. Top view of 10-lead model.
Tool ∆H
L
R −∆ H R
θ
5. Ten-lead finite element analysis and results The co-planarity between the bonding tool and the stage is a critical element for reliable thermocompression
L : ChipLength ∆ H : ToolPlanarity
θ : CentralAngle
Fig. 18. Sketch of concave surface definition.
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D.S. Liu et al. / Microelectronics Reliability 43 (2003) 935–943 2.5
Sinking Value(um)
Simulation data Experiment Data
2
1.5
1
0.2
0.4
0.6
0.8
1
Chip Length(1/L) Fig. 19. Comparison of simulation data with experimental data of gold bump sinking value by concave surface tool.
value experimental data. The simulation results show a similar phenomenon on the bump sink value. The approximate concave surface is therefore reasonable. The limited testing data may lead to the slight discrepancy between experimental and simulation results. 5.1. Tool profile effects Five concave tool surfaces were defined as shown in Fig. 20, to study the tool profile effects. Fig. 21 shows the gold bump sink value with a fixed bonding force, 30 gf. The sink value distributions were conjugated with convex tool surfaces. To achieve the minimum sink
Fig. 21. Gold bump sinking value, depending on tool planarity.
value requirement (1 lm), this study showed that the permissible tool tip surface variance should not exceed 1.25 lm for 30 gf bonding force. 5.2. Planarity angle effects The bonding tool was rotated using a small incline angle with respect to the pad surface to study the planarity angle effects. Fig. 22 shows the bump sink values with various planarity angles. Fig. 23 shows the bump deformation with various planarity angles and concave bonding surfaces (DH ¼ 1 lm). By coupling the
5 4.5 0o 0.005 o o 0.01 o 0.015 0.02o
Sinking Value(um)
4 3.5 3 2.5 2 1.5 1 0.5 0
0
0.2
0.4
0.6
0.8
1
Chip Length(1/L)
Fig. 20. Sketch of concave tool surface.
Fig. 22. Gold bump sinking value, depending on tool planarity angle.
D.S. Liu et al. / Microelectronics Reliability 43 (2003) 935–943
Acknowledgements
5
This research was supported by the R.O.C. NSF Foundation Grant NSC-91-2212-E-194-012 to the National Chung Cheng University.
0o 0.005o 0.01 o o 0.015 o 0.02
4
Sinking Value (um)
943
3
References 2
1
0
0
0.2
0.4
0.6
0.8
1
Chip Length(1/L) Fig. 23. Gold bump sinking value, depending on tool planarity angle with tool planarity (DH ¼ 0:75 lm).
planarity and flatness effects, the sink value distribution on chip is not a straight line. The acceptable incline angle for achieving the minimum sink value requirement is 0.005.
6. Conclusions This paper presents a three-dimensional FE computational model for simulating the TAB/ILB process using the explicit FE method. A new constitutive equation was developed to describe the stress–stain relationship of the copper beam lead at temperature ranges from 25 up to 250 C. Both single-lead and 10-lead models were developed and validated by comparison with the bump deformation obtained from actual TAB/ILB test samples. FE analysis was performed to examine the influences of tool position, tip profile and planarity angle on the bump deformation bonding process. The results obtained from the one-lead model demonstrate that a shorter beam lead is more sensitive to the tool position and increasing the beam lead length can decrease the effective plastic strain. The results obtained from the 10lead model showed that tool planarity and flatness have significant effects upon the bump sink value. The permissible tool tip surface variance should not exceed 1.25 lm, and the acceptable tool planarity angle is 0.005 for a fixed bonding force of 30 gf to achieve the minimum sink value requirement. The sink value distribution on the chip is not linear when the planarity and flatness effects are coupled.
[1] Lau JH. Handbook of tape automated bonding. New York: Van Nostrand Reinhold; 1992. [2] Lau JH, Rice DW, Harkins CG. Thermal stress analysis of tape automated bonding packages and interconnections. IEEE Trans Comp Hybrids Manufact Technol 1990; 13(1):182–7. [3] Zakel E, Reich H. Investigation of failure mechanisms of TAB-bonded chips during thermal aging. IEEE Trans Comp Hybrids Manufact Technol 1990;13(4):856–64. [4] Alpern P, Tilgner R. IR microscopic observation of plastic coated TAB inner lead bonds degrading during thermal cycling. IEEE Trans Comp Hybrids Manufact Technol 1992;15(1):114–7. [5] Tung CH, Kuo YS, Chang SM. Tape automated bonding inner lead bonded devices (TAB/ILB) failure analysis. IEEE Trans Comp Hybrids Manufact Technol 1993; 16(2):304–10. [6] Jog MA, Cohen IM, Ayyaswamy PS. Analysis and simulation of thermal transients and resultant stresses and strains in TAB packaging. J Electron Packaging Trans ASME 1993;115(1):34–8. [7] Cluff KD. Analysis of TAB inner lead fatigue in thermal cycle environments. IEEE Trans Comp Packaging Manufact Technol A 1995;18(1):101–7. [8] Atsumi K, Kashima N, Maehara Y, Mitsuhashi T, Komatsu T, Ochiai N. Inner lead bonding technique for 500-lead dies having a 90-lm lead pitch. IEEE Trans Comp Hybrids Manufact Technol 1990;13(1):222–8. [9] Lai JKL, Loo MC, Cheng KY. Effects of bond temperature and pressure on microstructures of tape automated bonding (TAB) inner lead bonds (ILB) with thin tape metallization. In: Electronic Components Technololgy Conference Proceedings, 1995. p. 819–26. [10] Oshige K, Nakanishi K. Development of a physical simulation system of ILB process. In: ASME Conference on Electronic Packaging Proceedings, 1992. p. 501–7. [11] Jung H, Lee T, Back J, Kim H. Optimization of TAB inner lead bonding process. In: 20th Annual IEEE SEMITHERM Semiconductor. Thermal Measurement Management Symposium Proceedings, 1996. p. 129–35. [12] Takahashi Y, Inoue M, Inoue K. Numerical analysis of fine lead bonding––effect of pad thickness on interfacial deformation. IEEE Trans Comp Packag Technol 1999; 22(2):291–8. [13] Liu DS, Chao YC. Effects of dopant, temperature and strain rate on the mechanical properties of micrometer gold bonding wire. TMS J Electron Mater 2003;32(3):159–65.
Numerical study on the bonding tool position, tip profile and planarity angle influences on TAB/ILB interconnection reliability D.S. Liu
a,*
, Y.C. Chao a, C.H. Lin a, G.S. Shen b, H.S. Liu
b
a
b
Department of Mechanical Engineering, National Chung Cheng University, 160, San-Hsing, Ming-Hsiung, Chia-Yi 621, Taiwan, ROC ChipMOS Technologies Inc., No. 5, NAN-KO Road, 7, Science-Bases Industrial Park, Tainan 744, Taiwan, ROC Received 23 September 2002; received in revised form 5 March 2003
Abstract Tape automated bonding (TAB) is a widely used interconnection technology for high-pincount and fine-pitch IC packaging. In this study, a three-dimensional computational model was developed for analyzing TAB inner lead bonding (ILB) process. This experimental study on the thermomechanical properties of copper leads was achieved using high precision micro-force tensile tests. A stress–stain relation between the copper lead and different temperature ranges was successfully implemented into the finite element model to study large plastic deformation in ILB formation. The resulting ILB lead profile and bump sinking values obtained from the simulations agreed well with the experimental observations from actual manufacturing data with the same bonding parameters. The tool position and lead length effects are analyzed to study the residual stress distribution after ILB. A 10-lead model was developed to study how the tool tip profile and planarity Ôangle affect the co-planarity between the bonding tool and the stage. The numerical results show that the permissible tool profile variance should not exceed 1.25 lm and the acceptable planarity angle is 0.005 to achieve the minimum bump deformation requirement. 2003 Elsevier Science Ltd. All rights reserved.
1. Introduction Tape automated bonding (TAB) technology is a momentous interconnection method in first level packaging for portable and handheld electrical production. TAB offers many advantages such as finer pitch and high I/O, lower profile, more precise geometry and gang bonding. This assembly technology corresponds well with the current trend in electronics [1]. The first assembly step in the TAB process involves bonding silicon chips to patterned metal-on-polymer tape, called inner lead bonding (ILB). The objective of ILB is to form a strong metallurgical bond between the gold bump on the
*
Corresponding author. Tel.: +886-5272-0411x33305; fax: +886-5272-0589. E-mail address: [email protected] (D.S. Liu).
die pad and the tin plated beam lead on the carrier tape (see Fig. 1). This process does not compromise the quality and electrical reliability of the final component. Thermocompression (T/C) gang bonding is the most commonly used TAB/ILB technology. This method can simultaneously bond all TAB beam leads to gold bumps on a chip using a bonding tool (thermode, see Fig. 2) with proper temperature, pressure and dwell time. This bonding method does increase throughput, but on chips with high-density I/O, planarity problems can lead to missing bonds, cracked chips or inconsistent pressure. Improper tool temperature, bonding force and tool position can cause fractures in the beam lead span near the bond pad (see Fig. 3). Currently, the packaging industry relies heavily on the use of tests to adjust the bonding parameters to meet the design requirements. This trialand-error process is time consuming and cost ineffective. Therefore, it is important to fully understand the
0026-2714/03/$ - see front matter 2003 Elsevier Science Ltd. All rights reserved. doi:10.1016/S0026-2714(03)00095-7
936
D.S. Liu et al. / Microelectronics Reliability 43 (2003) 935–943
Fig. 1. Schematic of ILB.
Fig. 2. Thermocompression bonding tool.
Fig. 3. Inner lead fracture.
dynamic characteristics of the TAB/ILB to develop better design tools that can quickly obtain optimum manufacturing parameters for various types of TAB packaging. In the past, much research has been performed to examine the possible TAB packaging failure modes due to thermal mismatch or cyclic thermal loading [2–7]. Few research studies attempted to investigate how the bonding parameters affect bonding strength and reliability. In experimental work, Atsumi et al. [8] studied the acceptable bonding conditions and the tool tip configuration for both the 182-lead and 504-lead test dies. Optimal bonding ranges for tool temperature, bonding pressure, and bonding time were found by examining the pull test bond strength. Lai et al. [9] investigated how bonding temperature and bonding pressure affected the microstructure of ILB joints. Oshige and Nakanishi [10] applied a simple two-dimensional numerical model to simulate the contact between a lead and bump pair. Only part of the lead close to the bump was modeled. Numerical results such as the deformation shape and strain distribution for the lead and bump were presented. Jung et al. [11] used the finite volume method to calculate the temperature distribution for each component during ILB, and concluded that the temperature of the heat cartridge in the bonding tool should not be above 520 C to avoid thermal damage (delamination) in the adhesive layer. Takahashi et al. [12] employed twodimensional contact finite element (FE) analysis to study the size ratio effect between the pad thickness and lead height in the ILB process. They found that if the pad thickness was decreased, the stress exhibited a peak under the pad that might cause damage to the die. Up to now, no research has presented a numerical model to examine how the geometric parameters such as the tool compression location affect the stress distribution on the beam lead span. Most studies used only a two-dimensional simple FE model to perform the analysis. This approach cannot be used to discuss the tool planarity and flatness effects. Therefore, in this research the TAB/ILB formation phenomena during the T/C bonding process was simulated using a single-lead threedimensional computational model and explicit FE analysis. Sn-plated Cu lead thin-strip samples were used as test specimens to perform micro-force tensile tests. The elasto-plastic stress–strain relations for the beam lead at various temperature ranges were combined into nonlinear dynamic FE code LS-DYNA3D to perform TAB/ILB formation simulations. The numerical results were used to study lead deformation, residual stress at the lead span after bonding and gold bump-downward length. A 10-lead FE model was also developed based on a single-lead model to study the effect of the tool tip profile and the tool and die coplanar surface. Both FE models were verified using actual TAB/ILB tests. The numerical simulation model presented here can provide
D.S. Liu et al. / Microelectronics Reliability 43 (2003) 935–943
better understanding of the fundamental formation phenomena during the ILB process and serve as a useful tool for better adjusting the bonding tool parameters to achieve the reliability requirements.
937
450
R.T. 400
100 oC 350 o
150 C
2. Copper lead constitutive model To determine the appropriate elasto-plastic constitutive model to quantify the plastic flow characteristic effects on the Cu lead in the ILB process, thermomechanical micro-force tensile testing was developed. The test sample material investigated in this study was electrodeposited copper film used as the raw material for Snplated Cu beam lead. Experiments were performed on an MTS-Tytron micro-force tester equipped with a 125 N load cell. An environmental chamber was attached to the testing machine to conduct thermal testing. The experimental setup is shown in Fig. 4. Copper films were cut into miniature thin-strip specimens with 40 mm · 5 mm (length by width) dimensions with 18 lm uniform thickness. To obtain consistent and accurate test results, the grip force was suitably adjusted to prevent the film from slipping or breaking. Specimen alignment was performed carefully to prevent warpage. An additional multi-axis translation stage and CCD camera was therefore used to provide the required motion and focus vision to ensure accurate alignment. Fig. 5 shows the tensile stress–strain curves at various temperature ranges. All of these curves appear the same where strain hardening occurs through the initial 0.2–2% strain and is followed by continued flow at minor stress increments. Examination of each curve in Fig. 5 illustrates that the strength decreases as the imposed temperature is increased. The maximum stress at room
Stress(MPa)
300 250
200 oC
200
250 oC 150 100
Copper Film Tensile Test
50 0
0
1
2
3
4
5
6
7
Strain(%)
Fig. 5. Stress–strain curves for copper film at strain rate 1 min1 .
temperature (440 MPa) is dramatically reduced to 150 MPa with the temperature increased to 250 C. An elastoplastic constitutive model with exponentiallaw form was proposed by the authors to describe the behavior of gold bonding wire [13]. The model was modified to adequately describe the behavior of the copper beam lead without strain hardening effects. The constitutive relations are written as follows: r ¼ E0 ð1 eae Þ E0 ðT Þ ¼ 481:3061 1:1775T 0:00046946T 2
ð1Þ T unit: C ð2Þ
aðT Þ ¼ 125:677 0:1331T þ 0:000694T 2
T unit: C ð3Þ
where E0 , a are the material constants. The first term in Eq. (1) expresses the elastic and transition portion of the stress–strain curve. The second term in Eq. (1) denotes the stress–strain relationship in a fully plastic range. A comparison of the predictions from Eq. (1) is presented in Fig. 6. The results calculated from the constitutive model agree well with the experimental data. 3. Finite element modeling and validation
Fig. 4. Schematic of experimental setup for copper film tensile test.
The analysis model configuration shown in Fig. 7 is based on the actual TAB/ILB process. Because of the symmetrical nature of the problem in hand, only half of the ILB model was used for the simulation. The dimensions of the gold bump were 100 lm · 40 lm · 17
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D.S. Liu et al. / Microelectronics Reliability 43 (2003) 935–943 500 450
R.T.
400
o
100 C
Stress(MPa)
350 o
150 C
300 250
o
200 C 200
250 oC
150
Copper Film Tensile Test
100
Model Prediction
50 0
Experimental data
0
1
2
3
4
5
6
7
Strain(%)
Fig. 6. Comparison of model prediction with experimental data of stress–strain curve of copper film.
lm. The thickness of the copper lead was 18 lm. The length of the copper lead beam was 274 lm. The FE model for analyzing the single lead TAB/ILB process is shown in Fig. 8. It consists of five major parts: the gold bump, the polyimide, the copper lead, the adhesive and the bonding tool. The FE mesh was created with a combination of 3-D element types. The gold bump, polyimide, copper lead, and adhesive were modeled as solid elements and the tool with rigid shell elements. The number of total solid elements is 4752, and 360 for the shell elements. The bonding tool temperature and dwell time during the ILB process is a crucial factor in determining the temperature distribution for each part of the ILB simulation model. Jung et al. [11] used the finite volume method to calculate the temperature distribution at various TAB tape areas with varying tool temperatures
and dwell time. In our actual TAB/ILB test, the bonding tool temperature was set at 500 C. The dwell time was set at 1.0 s. According to Ref. [11], the gold bump and lead were heated to about 220 C. Fig. 9 shows the stress–strain curve for the copper lead obtained from above and the gold bump from Ref. [13]. The material properties of the other parts are listed in Table 1. The FE code LS-DYNA3D was used in this study to simulate TAB/ILB transient formation process. In LSDYNA3D, the transient analysis is performed using an explicit direct-time-integration procedure and thereby avoids the need for matrix evaluation, assembly and decomposition at each time step as required by many implicit time-integration algorithms. The updated Lagrangian (UL) FE formulation adopted in this code accounts for both the material and geometric nonlinearity. Using this formulation the FE mesh is continuously updated during the calculation, and all variables calculated from UL FE are referred to the last calculated configuration. Piecewise linear elasto-plastic material model was used to input the stress–stain curves for copper lead and gold bump. The other parts of the TAB were modeled by elastic material model. Loading curve then input as a mechanical loading to simulate bonding tool pressure applied to inner lead and gold bump. The bonding tool speed was 5 mm/s downward. Some special techniques for TAB/ILB process FE simulation, such as the contact interfaces, must be carefully defined between tool and lead and an element-tied criterion based on the maximum contact force between lead and bump, are proposed and implemented. Once the contact force between the lead and bump reached the bonding pressure limit for a single lead element on the bottom surface of the lead, the top elements on the bump are tied. To validate this FE model, we simulated the TAB/ ILB process with various bonding pressures ranging from 20 to 45 gf and compared the lead profiles and gold bump sinking values with actual ILB test data. Fig. 10 compares the ILB lead deformation profile with a
R40
274 18
33
12
17
434 100
130
Fig. 7. Configuration of single lead model (in lm).
40
120
75
70
D.S. Liu et al. / Microelectronics Reliability 43 (2003) 935–943
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Fig. 8. Single lead FE model: (a) oblique view and (b) side view.
250
Table 1 Material properties of single lead model
Stress(MPa)
200
Copper Lead Gold Bump
Adhesive Polyimide Tool
150
Elastic modulus (GPa)
PoissonÕs ratio
0.6367 3.15 Rigid
0.28 0.3
Yield stress (GPa)
Plastic hardening modulus (GPa)
0.085
0.97
100
50
0
0
1
2
3
4
5
6
7
8
9
Strain(%)
Fig. 9. Stress–strain curve of copper film and gold bump.
bonding force test results equal to 30 gf. Fig. 11 shows a schematic of the bump sink value definition and compares the bump sink values with the test results at various bonding pressures. Both agree very well with the test results.
Fig. 10. Comparison of the deformed profile of copper lead (upper: test results, lower: numerical results).
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D.S. Liu et al. / Microelectronics Reliability 43 (2003) 935–943 3.5 Experiment Simulation
Sinking Value(µm)
3
2.5
2 A
C
B
1.5
1
0.5
Fig. 13. Maximum Von-Mises stress.
A : L ead Thickness B : Bump Height C : Measurement Sinking Value = (A-B)-C
20
30
40
50
Bonding Force(gf)
Fig. 11. Comparison of numerical data with experimental data of gold bump sinking value.
4. Single lead analysis and results 4.1. Failure mode of beam lead and bonding force effects
Fig. 14. SEM micrograph of fracture of a beam lead.
240 220
Max. Von Mises Stress(MPa)
From the simulation the TAB/ILB process might be separated into four steps, as shown in Fig. 12. Step one, the tool makes initial contact with the end of the lead beam. Step two, the lead slides along the tool tip and makes contact with the bump-top surface. The sliding distance is about 50 lm. Step three, the tool compresses the lead top surface and ties the lead bottom surface with the bump. Step four, the tool rises up after completing the ILB process. Fig. 13 shows the effective (VonMises) stress contour distribution of the deformed beam lead after the third step. The maximum effective stress value occurred at the elements close to the tool tip. This observation is consistent with the failure mode from actual TAB/ILB tests, as shown in Fig. 14. Fig. 15 shows the maximum effective stresses vs. various bonding forces at the beam lead near the tool tip and in the gold bump. A larger bonding force increases
200 Inner Lead Gold Bump
180 160 140 120 100 80 60 40 20 20
25
30
35
40
45
Bonding Force(gf)
Fig. 15. Max effective Stress on lead and gold bump.
Fig. 12. ILB process.
the effective stress in the gold bump and copper lead. Excessive bonding force might be considered to be one of the driving forces causing damage to the silicon chip. Conversely, insufficient bonding force decreases the bump deformation and might produce a poor mechan-
D.S. Liu et al. / Microelectronics Reliability 43 (2003) 935–943
Effective Plastic Strain(%)
6
Lead Length 0.274mm Lead Length 0.324mm Lead Length 0.374mm
5
4
3
2 -20
0
20
40
60
Tool Position(um)
Fig. 16. Effective plastic strain on tool lip lead.
ical or electric bond. In actual TAB/ILB tests, the bonding tool force was set at 30 gf. In the following parameter study all results were calculated under bonding force equal to 30 gf. The bump deformation was the criteria for good bonding. According to the manufacturerÕs pull test data, a bump deformation range from 1 to 7 lm represents a good connection.
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gang bonding, especially as chip sizes become larger and the bump and lead geometry become smaller. In addition to the bonding tool planarity setup, the profile of the bonding tip surface at various elevated temperature is also critical to a successful gang ILB process. As the temperature increases up to 500 C, the profile of the tool tip surface exhibits a concave shape with approximately 1 lm of deformation at the periphery of the tip. A 10-lead FE model was developed to study the tool profile variation and planarity problems in the ILB process. The numerical results were compared with the experimental data. Fig. 17 shows the 10-lead FE model. The distance between the two leads is 50 lm. The material properties are the same as that for a single lead model. The tool surface was approximated as a concave circular arc that could be defines as long as the tool planarity distance (DH ) and chip length are given (see Fig. 18). To verify the accuracy of the computational model, simulations were carried out with DH ¼ 1 lm and chip length equal 16.85 mm according to the actual TAB/ ILB bonding experiments. The actual bonding samples were bonded using a Shibaura TAB bonder. The bump deformation for each sample was measured using a scanning electron microscope (SEM). Each experimental result was averaged from five testing specimens. Fig. 19 compares the numerical results with the gold bump sink
4.2. Bonding tool position / lead length effects The original horizontal distance between the tool tip and polyimide was 274 lm. This was chosen as the base line position (zero position). Four additional tool positions were examined; positions 1 to 3 involved moving the bonding tool horizontally with respect to the zero position 20, 40, and 60 lm toward to the edge of the polyimide. Position 4 involved moving the tool inversely 20 lm ()20 lm position) with respect to the original tool position. The single-lead FE model was performed to analyze these four cases to understand how the tool position affected the residual stress/strain. In addition, the length of the lead beam was increased to 324 and 374 lm to study the lead length effect. Fig. 16 illustrates the effective plastic strain on the lead near the tool tip. Increasing the beam lead length lead to a decrease in the effective plastic strain. Moreover, the slope of the curves in Fig. 16 decreased with increasing beam lead length, indicating that a shorter beam lead is more sensitive to the tool position.
Fig. 17. Top view of 10-lead model.
Tool ∆H
L
R −∆ H R
θ
5. Ten-lead finite element analysis and results The co-planarity between the bonding tool and the stage is a critical element for reliable thermocompression
L : ChipLength ∆ H : ToolPlanarity
θ : CentralAngle
Fig. 18. Sketch of concave surface definition.
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D.S. Liu et al. / Microelectronics Reliability 43 (2003) 935–943 2.5
Sinking Value(um)
Simulation data Experiment Data
2
1.5
1
0.2
0.4
0.6
0.8
1
Chip Length(1/L) Fig. 19. Comparison of simulation data with experimental data of gold bump sinking value by concave surface tool.
value experimental data. The simulation results show a similar phenomenon on the bump sink value. The approximate concave surface is therefore reasonable. The limited testing data may lead to the slight discrepancy between experimental and simulation results. 5.1. Tool profile effects Five concave tool surfaces were defined as shown in Fig. 20, to study the tool profile effects. Fig. 21 shows the gold bump sink value with a fixed bonding force, 30 gf. The sink value distributions were conjugated with convex tool surfaces. To achieve the minimum sink
Fig. 21. Gold bump sinking value, depending on tool planarity.
value requirement (1 lm), this study showed that the permissible tool tip surface variance should not exceed 1.25 lm for 30 gf bonding force. 5.2. Planarity angle effects The bonding tool was rotated using a small incline angle with respect to the pad surface to study the planarity angle effects. Fig. 22 shows the bump sink values with various planarity angles. Fig. 23 shows the bump deformation with various planarity angles and concave bonding surfaces (DH ¼ 1 lm). By coupling the
5 4.5 0o 0.005 o o 0.01 o 0.015 0.02o
Sinking Value(um)
4 3.5 3 2.5 2 1.5 1 0.5 0
0
0.2
0.4
0.6
0.8
1
Chip Length(1/L)
Fig. 20. Sketch of concave tool surface.
Fig. 22. Gold bump sinking value, depending on tool planarity angle.
D.S. Liu et al. / Microelectronics Reliability 43 (2003) 935–943
Acknowledgements
5
This research was supported by the R.O.C. NSF Foundation Grant NSC-91-2212-E-194-012 to the National Chung Cheng University.
0o 0.005o 0.01 o o 0.015 o 0.02
4
Sinking Value (um)
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3
References 2
1
0
0
0.2
0.4
0.6
0.8
1
Chip Length(1/L) Fig. 23. Gold bump sinking value, depending on tool planarity angle with tool planarity (DH ¼ 0:75 lm).
planarity and flatness effects, the sink value distribution on chip is not a straight line. The acceptable incline angle for achieving the minimum sink value requirement is 0.005.
6. Conclusions This paper presents a three-dimensional FE computational model for simulating the TAB/ILB process using the explicit FE method. A new constitutive equation was developed to describe the stress–stain relationship of the copper beam lead at temperature ranges from 25 up to 250 C. Both single-lead and 10-lead models were developed and validated by comparison with the bump deformation obtained from actual TAB/ILB test samples. FE analysis was performed to examine the influences of tool position, tip profile and planarity angle on the bump deformation bonding process. The results obtained from the one-lead model demonstrate that a shorter beam lead is more sensitive to the tool position and increasing the beam lead length can decrease the effective plastic strain. The results obtained from the 10lead model showed that tool planarity and flatness have significant effects upon the bump sink value. The permissible tool tip surface variance should not exceed 1.25 lm, and the acceptable tool planarity angle is 0.005 for a fixed bonding force of 30 gf to achieve the minimum sink value requirement. The sink value distribution on the chip is not linear when the planarity and flatness effects are coupled.
[1] Lau JH. Handbook of tape automated bonding. New York: Van Nostrand Reinhold; 1992. [2] Lau JH, Rice DW, Harkins CG. Thermal stress analysis of tape automated bonding packages and interconnections. IEEE Trans Comp Hybrids Manufact Technol 1990; 13(1):182–7. [3] Zakel E, Reich H. Investigation of failure mechanisms of TAB-bonded chips during thermal aging. IEEE Trans Comp Hybrids Manufact Technol 1990;13(4):856–64. [4] Alpern P, Tilgner R. IR microscopic observation of plastic coated TAB inner lead bonds degrading during thermal cycling. IEEE Trans Comp Hybrids Manufact Technol 1992;15(1):114–7. [5] Tung CH, Kuo YS, Chang SM. Tape automated bonding inner lead bonded devices (TAB/ILB) failure analysis. IEEE Trans Comp Hybrids Manufact Technol 1993; 16(2):304–10. [6] Jog MA, Cohen IM, Ayyaswamy PS. Analysis and simulation of thermal transients and resultant stresses and strains in TAB packaging. J Electron Packaging Trans ASME 1993;115(1):34–8. [7] Cluff KD. Analysis of TAB inner lead fatigue in thermal cycle environments. IEEE Trans Comp Packaging Manufact Technol A 1995;18(1):101–7. [8] Atsumi K, Kashima N, Maehara Y, Mitsuhashi T, Komatsu T, Ochiai N. Inner lead bonding technique for 500-lead dies having a 90-lm lead pitch. IEEE Trans Comp Hybrids Manufact Technol 1990;13(1):222–8. [9] Lai JKL, Loo MC, Cheng KY. Effects of bond temperature and pressure on microstructures of tape automated bonding (TAB) inner lead bonds (ILB) with thin tape metallization. In: Electronic Components Technololgy Conference Proceedings, 1995. p. 819–26. [10] Oshige K, Nakanishi K. Development of a physical simulation system of ILB process. In: ASME Conference on Electronic Packaging Proceedings, 1992. p. 501–7. [11] Jung H, Lee T, Back J, Kim H. Optimization of TAB inner lead bonding process. In: 20th Annual IEEE SEMITHERM Semiconductor. Thermal Measurement Management Symposium Proceedings, 1996. p. 129–35. [12] Takahashi Y, Inoue M, Inoue K. Numerical analysis of fine lead bonding––effect of pad thickness on interfacial deformation. IEEE Trans Comp Packag Technol 1999; 22(2):291–8. [13] Liu DS, Chao YC. Effects of dopant, temperature and strain rate on the mechanical properties of micrometer gold bonding wire. TMS J Electron Mater 2003;32(3):159–65.