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Characterisation of constitutive behaviour of SnAg, SnAgCu and SnPb solder in flip chip joints
Sensors and Actuators A 99 (2002) 188±193
Characterisation of constitutive behaviour of SnAg, SnAgCu and SnPb solder in ¯ip chip joints S. Wiese*, F. Feustel, E. Meusel Technische UniversitaÈt Dresden, Institut fuÈr Halbleiter-und Mikrosystemtechnik, 01062 Dresden, Germany
Abstract The constitutive behaviour of Sn96.5Ag3.5, Sn95.5Ag4Cu0.5 and Sn63Pb37 solder was investigated on ultra small ¯ip chip solder joints (V 1 10 12 m3). In order to run experiments on these small specimens, a micro shear tester has been designed. The tester is optimised to achieve high precision. It is able to record force displacement hysteresis with a resolution of better than 1 mN and 20 nm for force and displacement measurements, respectively. It was found that the creep behaviour of the lead-free solders exhibit a very high dependence on stress (Sn96.5Ag3.5: n 11; Sn95.5Ag4Cu0.5: n 18), while the eutectic SnPb showed a low stress dependence of n 2. # 2002 Elsevier Science B.V. All rights reserved. Keywords: SnAg; SnAgCu; SnPb; Constitutive behaviour; Lead-free solder; Flip chip
1. Introduction For a long time, the eutectic SnPb solder is being used as the standard material for joining of electronic components because of its suitable physical properties and low cost. Lead, however, need to be totally removed from industrial products in the coming years. Among the diversity of alternative lead-free solders, the SnAg/SnAgCu system seems to make the race. However, at the moment of time, the mechanical behaviour of SnAg/SnAgCu eutectic solders is not well understood. The rare data that exists for the SnAg/ SnAgCu system shows enormous scatter. This does not allow to make reliable predictions by FEM simulations about the thermomechanical fatigue of solder joints in electronic packages.
and applied stress. e_ f T; s
If a low or medium stress is applied the stress dependence steady-state creep rate of most metals and alloys can be described by s n e_ A1 (2) E where n has usually a value between 3 and 10 and is called stress exponent. At high stresses, this simple power law creep behaviour breaks down and the creep rate (_e) increases more quickly with the applied stress (s). Therefore, it is called power law breakdown creep behaviour and can be described by the following empirical relation e_ A2 exp B2 s
2. Theory of creep deformation The creep of metals depends on a large number of variables and thus dif®cult to describe. Therefore, in many technological applications only the description of the steady-state creep behaviour is used. A steady-state creep is reached after a constant stress has been applied and the material has already passed through a transient phase of creep. The steady-state creep rate is a function of temperature *
Corresponding author. Tel.: 49-351-463-3172; fax: 49-351-463-7172. E-mail address: [email protected] (S. Wiese).
(1)
(3)
In order to ®nd a closed formulation for power law creep (low and medium stresses) and power law breakdown (high stresses) the following equation has been widely established B3 s n e_ A3 sinh (4) E Since creep involves many thermally activated processes and mechanisms, the temperature dependence of steadystate creep rate can usually be expressed by the Arrhenius law Q _e e_ 0 exp (5) kT
0924-4247/02/$ ± see front matter # 2002 Elsevier Science B.V. All rights reserved. PII: S 0 9 2 4 - 4 2 4 7 ( 0 1 ) 0 0 8 8 0 - 9
S. Wiese et al. / Sensors and Actuators A 99 (2002) 188±193
189
where e_ is a pre-factor that accounts for the creep rate dependence on stress, k the Boltzmann's constant and Q the activation energy of the dominant creep mechanism. Q can obtain different values in dependence of temperature and stress.
the steady-state strain rate. Nevertheless, the introduction of sth is useful because as experiments have shown the value of correlates with volume fraction as well as with the size and shape of the dispersed hard particles [3,4]
2.1. Superplasticity
3. Experimental
When a metal is pulled in tension, it usually breaks after a fairly small elongation. However, Pearson [1] reported that eutectic SnPb can be pulled to elongations up to 2000%. For the eutectic SnPb was found [2] that the creep behaviour at low and medium stresses can be divided into three distinct regions: region I (low strain rates) Nabarro±Heering creep; region II (intermediate strain rates) superplastic behaviour; region III (high strain rates) climb controlled dislocation creep. Although experimental results for region I are contradictory, there is agreement about the creep behaviour in regions II and III, which is usually expressed by double power law equation: snII QII QIII nIII e_ AII p exp AIII s exp (6) d kT kT
The specimen (Fig. 1) consisted of two silicon chips, which were connected by four ¯ip chip joints (each on one corner) to each other. SnAg and SnAgCu ¯ip chip bumps were printed 200 mm 200 mm Cu pads. Sn63Pb37 ¯ip chip bumps were deposited by electroplating on 100 mm 100 mm Cu pads. The joint shape was chosen hyperbolic. In contrast to the usual barrel shaped joints, in hyperbolic shaped joints, all strains and stresses will concentrate in the centre away from the interface to the under bump metalisation. That way, any interference with the interfacial region to the under bump metalisation is avoided. One of the problems that comes along which such type of real ¯ip chip specimen is to design a matching test setup for mechanical experiments. The imposed loading conditions should be reversal shear under isothermal conditions with appropriate cyclic strain rates. This requires a well controlled movement of the two silicon chips of the specimen against each other in such a way that all bending moments and out of plane forces are eliminated or at least minimised. This is realised through a symmetric grip con®guration of two identical specimens (Fig. 2). The test setup is driven by a piezoelectric translator, which provides a smooth movement with sub-nanometer resolution for a wide range of velocities. A force sensor is mounted between the inner grip and the
where d is the grain size, p the grain size exponent and AII, AIII the pre-factors. In region II, the SnPb eutectic has usually a low stress dependence with nII 2, when it exhibits superplastic deformation. 2.2. Creep in precipitation strengthened alloys The creep strength of two-phase or multiphase alloys can be dramatically higher then this of single phase alloys because of interactions between dislocations and metallurgical obstacles. Precipitated particles act as the most effective one. Theoretical models of dislocation movement in particle strengthened metals assume, that at high temperatures dislocations can undergo non-planar motion by climb, which allows the dislocation segment arrested at a particle bulge out of the slip plane and ®nally surmount the particle [3]. Since the traditional models of dislocation movement in metals were unable to explain high values of stress exponent n (upto 40) or unrealistic values of the apparent activation energy (Q), which were obtained from creep experiments on precipitation strengthened alloys, a creep strength increment sth s sm (s: applied stress; sm: creep strength of the referring particle-free metal matrix) was introduced to theory. With the empirical value of sth, the steady-state creep rate dependence on stress in precipitation strengthened alloys can be expressed as s s n th e_ A4 (7) E where n is the stress exponent for steady-state deformation of the reference material having the same composition as the matrix that the particles are dispersed in. The creep strength increment sth is no true threshold because it decreases with
Fig. 1. Flip chip specimens.
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S. Wiese et al. / Sensors and Actuators A 99 (2002) 188±193
Fig. 2. Test setup.
piezoelectric translator. The translation direction is in line with the gravity axis, in order to avoid any bearing. Bearings cause frictional forces of at least 150 mN, which will not allow a high resolution force measurement. A high resolution displacement measurement is achieved by the application of a dual beam laser interferometer. Each laser beam is re¯ected at one of two neighbouring edges of the specimen silicon chips. In order to minimise errors due to temperature changes or vibrations, the laser interferometer heads are borne symmetrically to the specimen. The achieved true precision of the microtester is 20 nm for displacement and 2 mN for force. A more detailed description of the test setup and the experimental programme is given in [5]. 4. Results The steady-state creep data of eutectic SnAg/SnAgCu solders is plotted versus shear stress in Fig. 3. All data was experimentally determined at temperatures between 5 and 50 8C. Fig. 4 compares the corresponding microstructure of the ¯ip chip solder (V 1 10 12 m3) and compares it with
Fig. 3. Creep behaviour of Sn96.5Ag3.5/Sn95.5Ag4Cu0.5 solders.
that of an Sn95.5Ag4Cu0.5 bulk specimen (V 1 10 8 m3). The volume ratio is 1:10,000. The microstructure is composed of Ag3Sn intermetallics (dark) distributed in a b-Sn matrix (light). Intermetallics of Z-Cu6Sn5 appear also light and therefore, cannot been distinguished from the b-Sn matrix. Both specimens were soldered using SMT similar re¯ow pro®le. Although the differences in microstructure between both specimens enormous, it is very dif®cult to say, if these differences derive from volume effects or from differences in solder composition. Since the alloying content of Ag (4%) and Cu (0.5%) is relatively little, the solution of UBM metals (Au, Ni, Cu, etc.) during bumping and soldering processes may change the composition of the solder and thus the solidi®cation process changes signi®cantly, because the molten solder has moved away from the eutectic in the complex ternary SnAgCu system. The steady-state displacement rate data, that was gained on ¯ip chip specimens of eutectic SnPb solder at temperatures of 5 and 50 8C is plotted versus shear force in Fig. 5. All odd numbers correspond to specimens with an as cast microstructure (grain size <1 mm) and all even numbers correspond to specimens with a coarsened microstructure (grain size 10 mm). Force and displacement rate values were not converted into stress and strain rate values because the material model calibration was performed through FEM simulation [5]. The microstructure of Sn63Pb37 in a bulk solder volume (V 1 10 8 m3) and in a ¯ip chip joint (V 1 10 12 m3) is shown as a back scattered electron images in Fig. 6. Both specimens were soldered with a SMT similar re¯ow pro®le. What is common on both microstructures are lead-rich (light) clusters of globular shape with a diameter (d) between 5 and 20 mm. The rest of microstructure appears different. While the ¯ip chip cross-section shows a very homogeneous distribution of globular shaped lead-rich phases (light) in a tin-rich (dark) matrix (mean phase intercept approximately 0.9 mm), the cross-section of the bulk solder contains a non-uniform distribution of phases. The mean phase intercept starts at approximately 4 mm at coarse structured regions and ends at approximately 0.5 mm at ®ne
S. Wiese et al. / Sensors and Actuators A 99 (2002) 188±193
191
Fig. 4. Optical micrograph of Sn95.5Ag4Cu0.5 solder in a flip chip and a bulk specimen.
undercooling (DT 55 K) to nucleate, the nucleation of lead in the presence of tin starts with only little undercooling (DT 0:25 K) [6]. Since the undercooling that is needed for the homogeneous nucleation of lead on primary lead-rich crystals is very similar to the undercooling required for the heterogeneous nucleation of lead on primary lead-rich crystals, it does not matter if the composition of the solder is slightly hyper- or hypoeutectic, when the temperature falls below the eutectic point lead-rich phases start to grow until the composition of the liquid phase has been shifted to 80 wt.% Sn±20 wt.% Pb. At this composition the temperatures at which tin and lead-rich phase nucleates are very nearly the same [7]. Fig. 5. Creep behaviour of Sn63Pb37 solder.
structured regions. This very speci®c microstructural appearance of SnPb solder can be explained by the nonreciprocal nucleation behaviour of the bimaterial tin±lead system. While tin in the presence of lead needs a high
5. Discussion In order to compare the steady-state creep behaviour of Sn63Pb37, Sn96.5Ag3.5 and Sn95.5Ag4Cu0.5, the two governing parameters n and Q of Eqs. (1)±(7) should be
Fig. 6. Back scattered electron images of the Sn63Pb37 solder in a bulk and flip chip specimen.
192
S. Wiese et al. / Sensors and Actuators A 99 (2002) 188±193 Table 2 Parameters for creep model (Eq. (9)) C1 (s 1) Sn96.5Ag3.5 Sn95.5Ag4Cu0.5
5 10 2 10
6 21
C2
C3 (MPa 1)
C4 (kJ/mol)
11 18
1 1
79.8 83.1
behaviour of precipitation strengthened materials is modelled by a simple power law C4 (9) e_ C1 C3 sC2 exp kT
Fig. 7. Comparison of creep of Sn96.5Ag3.5, Sn95.5Ag4Cu0.5 and Sn63Pb37 (all data points are corrected for T 300 K)
considered separately. For this reason the steady-state creep rates from the tests conducted at temperatures between 5 and 50 8C on ¯ip chip joints, were corrected for a temperature of T 27 8C using Eq. (5) and assuming a apparent activation energy. Fig. 7 shows the comparison of selected creep data of the three solders investigated. At room temperature Sn63Pb37 and Sn96.5Ag3.5 solder show nearly the same absolute creep rates at high stresses. At low stresses, however, Sn63Pb37 solder reduces its creep rate dependence on stress to an exponent of approximately n 2, whereas, a constant stress exponent of n 11 was observed for Sn96.5Ag3.5 solder. The low stress exponent of n 2 for Sn63Pb37 indicates that the solder in ¯ip chip joints deforms at least partially superplastically, when stresses are low. Modelling of creep behaviour of (near) eutectic tin lead solder (Table 1) was performed using the sinh formulation from Eq. (4) C4 C2 e_ C1 sinh C3 s exp (8) kT The high stress exponent of n 11 for Sn96.5Ag3.5 can be explained by a high value of sth (Eq. (7)). Since the cooling rate during solidi®cation is very high for a ¯ip chip joint, it is very likely that the solder in the ¯ip chip joint contains a high number of very small Ag3Sn precipitates. In contrast, the same solder in a bulk specimen that was solidi®ed at low cooling rate contains a smaller number of larger Ag3Sn precipitates (Fig. 4). At the same volume fraction of Ag3Sn a higher number of smaller particles acts as a more potent obstacle to the movement of dislocations than a smaller number of larger particles, which results in high values of for the solder in ¯ip chip joints. If creep Table 1 Parameters for creep model (Eq. (8))
Sn63Pb37
C1 (s 1)
C2
C3 (MPa 1)
C4 (kJ/mol)
10
2
0.2
44.9
the derivation of Eq. (7) gives ns @sth 1 C 2 n0 s sth @s T
(10)
Since n is a constant and is given by the matrix metal (stress exponent of b-Sn n 4 5), it can be concluded from Eq. (10), that n0 increases with increasing sth. Therefore, it can be assumed that the creep properties of precipitation strengthened material like Sn96.5Ag3.5 show a strong dependence on the cooling conditions that in¯uence the microstructure of the solder. Since it is very dif®cult to determine the value of sth experimentally it was preferred to model the creep behaviour of Sn96.5Ag3.5 solder (Table 2) with the simple power law from Eq. (9). Sn95.5Ag4Cu0.5 solder in ¯ip chip joints showed a stress exponent of n 18 (actually n0 18 for Eq. (10)) which is signi®cantly higher than that found on the Sn96.5Ag3.5 material. In addition, the ternary eutectic of the SnAgCu alloy showed much lower absolute creep rates than the binary eutectic of the SnAg material. This extreme increase in creep resistance may result from tiny (5 and 50 nm) precipitates of Z-Cu6Sn5 in the b-tin matrix [8,9]. Such small particles act as potent obstacles for dislocation movement and thus the creep resistance of an SnAgCu solder can be increased dramatically. For creep modelling of the Sn95.5Ag4Cu0.5 (Table 2) ¯ip chip solder the simple power law (Eq. (9)) as for the Sn96.5Ag3.5 was chosen. The apparent activation energy of creep Q (Eq. (5)) found on ¯ip chip specimens of Sn63Pb37 solder is in accordance with what is reported in literature. However, there are strong differences between literature values [10±14] of activation energy Q and these found in this study for Sn96.5Ag3.5 and Sn95.5Ag4Cu0.5 solder. These differences result supposedly from the fact that the ®t function (Eq. (9)) include implicit the creep resistance increment. Therefore, the parameter C4 (apparent activation energy) in the Eq. (9) corresponds to a corrected activation energy Q0 , that results from Eqs. (5) and (7) [4] n 1 dE n @sth Q0 Q kT 2 (11) E dT s sth @s s where Q is the apparent activation energy for the creep mechanisms in the precipitation strengthened material,
S. Wiese et al. / Sensors and Actuators A 99 (2002) 188±193
which is corrected by the temperature dependence of Young's modulus and sth. It is reasonable to assume that the temperature dependence of sth is similar to that of E. Since dE=dT < 0, Q0 increases when the difference s sth becomes smaller. This means larger values of n0 (C2), which result from higher correspond to larger values of Q0 (C4). 6. Conclusions The steady-state creep rate was determined by reversible constant load shear tests on Sn96.5Ag3.5, Sn95.5Ag4Cu0.5 and Sn63Pb37, ¯ip chip solder joints. The stress exponent found on Sn96.5Ag3.5 solder (n 11) and Sn95.5Ag4Cu0.5 solder (n 18) was signi®cantly higher than that found on Sn63Pb37 solder (n 2) in the low stress region. Also the apparent activation energy was higher for the Sn96.5Ag3.5 (Q 79:8 kJ/mol) and the Sn95.5Ag4Cu0.5 (Q 83:1 kJ/ mol) solder then for the Sn63Pb37 (Q 44:9 kJ/mol) solder. Microstructural analysis and theoretical assumptions turn out that the high values for n and Q that were found on the Sn96.5Ag3.5 and Sn95.5Ag4Cu0.5 solder might be caused by small precipitates of Ag3Sn and Z-Cu6Sn5 intermetallics that are ®nely dispersed in the b-tin matrix. References [1] C.E. Pearson, The viscous properties of extruded eutectic alloys of lead±tin and bismuth±tin, J. Inst. Metals 54 (1934) 111±124. [2] D.H. Avery, W.A. Backofen, Trans. Am. Soc. Metals 58 (1965) 551.
193
[3] B. Reppich, Particle strengthening, in: R.W. Cahn, P. Haasen, E.J. Kramer (Eds.), Materials Science and TechnologyÐA Comprehensive Treatment, Vol. 6, VCH, Weinheim, 1993, pp. 311±355. [4] J. Cadek, Creep in Metallic Materials, Elsevier, Amsterdam, 1988. [5] S. Wiese, F. Feustel, S. Rzepka, E. Meusel, Experimental characterisation of material properties of Sn63Pb37 flip chip solder joints, in: D.J. Belton, M. Gaynes, E.G. Jacobs, R. Pearson, T. Wu (Eds.), Proceedings of the Tenth MRS Symposium on Electronic Packaging Materials Science, Vol. 515, Materials Research Society, San Francisco, 14±16 April 1998, pp. 233±238. [6] B.E. Sundquist, L.F. Mondolfo, Heterogeneous nucleation in the liquid-to-solid transformation in alloys, Trans. Metallurg. Soc. AIME 221 (1961) 157±163. [7] J.H. Hollomon, D. Turnbull, Solidification of lead±tin alloy droplets, J. Metal Trans. AIME 3 (1951) 803±805. [8] L.E. Felton, K. Rajan, P.J. Ficalora, P. Singh, n-Cu6Sn5 precipitates in Cu/PbSn solder joints, Scripta Metallurgica et Materialia 25 (1991) 2329±2333. [9] L. Xiao, J. Liu, Z. Lai, L. Ye. A. ThoÈlen, Characterization of mechanical properties of bulk lead-free solders, in: Proceedings of the International Symposium on Advanced Packaging Materials, Braselton, GA, 6±8 March 2000, pp. 145±151. [10] E.I. Stromswold, Characterization of eutectic tin±silver solder joints, Dissertation, University of Rochester, 1993. [11] R. Darveaux, K. Banerji, Constitutive relations for tin-based solder joints. IEEE Transactions on CHMT 15 (1992) 1013±1024. [12] Z. Guo, Y.H. Pao, H. Conrad, Plastic deformation kinetics of Sn95.5Cu4Ag0.5 solder joints, J. Electron. Packaging 117 (1995) 100±104. [13] D.R. Frear, The mechanical behaviour of interconnect materials for electronic packaging, J. Metals 48 (1996) 49±53. [14] A. Schubert, H. Walter, A. Gollhardt, B. Michel, Materials mechanics and reliability issues of lead-free solder interconnects, in: Proceedings of the Second Conference on Benefiting from Thermal and Mechanical Simulation in Micro-ElectronicsÐEuroSimE 2001, Paris, 9±11 April 2001, pp. 171±178.
Characterisation of constitutive behaviour of SnAg, SnAgCu and SnPb solder in ¯ip chip joints S. Wiese*, F. Feustel, E. Meusel Technische UniversitaÈt Dresden, Institut fuÈr Halbleiter-und Mikrosystemtechnik, 01062 Dresden, Germany
Abstract The constitutive behaviour of Sn96.5Ag3.5, Sn95.5Ag4Cu0.5 and Sn63Pb37 solder was investigated on ultra small ¯ip chip solder joints (V 1 10 12 m3). In order to run experiments on these small specimens, a micro shear tester has been designed. The tester is optimised to achieve high precision. It is able to record force displacement hysteresis with a resolution of better than 1 mN and 20 nm for force and displacement measurements, respectively. It was found that the creep behaviour of the lead-free solders exhibit a very high dependence on stress (Sn96.5Ag3.5: n 11; Sn95.5Ag4Cu0.5: n 18), while the eutectic SnPb showed a low stress dependence of n 2. # 2002 Elsevier Science B.V. All rights reserved. Keywords: SnAg; SnAgCu; SnPb; Constitutive behaviour; Lead-free solder; Flip chip
1. Introduction For a long time, the eutectic SnPb solder is being used as the standard material for joining of electronic components because of its suitable physical properties and low cost. Lead, however, need to be totally removed from industrial products in the coming years. Among the diversity of alternative lead-free solders, the SnAg/SnAgCu system seems to make the race. However, at the moment of time, the mechanical behaviour of SnAg/SnAgCu eutectic solders is not well understood. The rare data that exists for the SnAg/ SnAgCu system shows enormous scatter. This does not allow to make reliable predictions by FEM simulations about the thermomechanical fatigue of solder joints in electronic packages.
and applied stress. e_ f T; s
If a low or medium stress is applied the stress dependence steady-state creep rate of most metals and alloys can be described by s n e_ A1 (2) E where n has usually a value between 3 and 10 and is called stress exponent. At high stresses, this simple power law creep behaviour breaks down and the creep rate (_e) increases more quickly with the applied stress (s). Therefore, it is called power law breakdown creep behaviour and can be described by the following empirical relation e_ A2 exp B2 s
2. Theory of creep deformation The creep of metals depends on a large number of variables and thus dif®cult to describe. Therefore, in many technological applications only the description of the steady-state creep behaviour is used. A steady-state creep is reached after a constant stress has been applied and the material has already passed through a transient phase of creep. The steady-state creep rate is a function of temperature *
Corresponding author. Tel.: 49-351-463-3172; fax: 49-351-463-7172. E-mail address: [email protected] (S. Wiese).
(1)
(3)
In order to ®nd a closed formulation for power law creep (low and medium stresses) and power law breakdown (high stresses) the following equation has been widely established B3 s n e_ A3 sinh (4) E Since creep involves many thermally activated processes and mechanisms, the temperature dependence of steadystate creep rate can usually be expressed by the Arrhenius law Q _e e_ 0 exp (5) kT
0924-4247/02/$ ± see front matter # 2002 Elsevier Science B.V. All rights reserved. PII: S 0 9 2 4 - 4 2 4 7 ( 0 1 ) 0 0 8 8 0 - 9
S. Wiese et al. / Sensors and Actuators A 99 (2002) 188±193
189
where e_ is a pre-factor that accounts for the creep rate dependence on stress, k the Boltzmann's constant and Q the activation energy of the dominant creep mechanism. Q can obtain different values in dependence of temperature and stress.
the steady-state strain rate. Nevertheless, the introduction of sth is useful because as experiments have shown the value of correlates with volume fraction as well as with the size and shape of the dispersed hard particles [3,4]
2.1. Superplasticity
3. Experimental
When a metal is pulled in tension, it usually breaks after a fairly small elongation. However, Pearson [1] reported that eutectic SnPb can be pulled to elongations up to 2000%. For the eutectic SnPb was found [2] that the creep behaviour at low and medium stresses can be divided into three distinct regions: region I (low strain rates) Nabarro±Heering creep; region II (intermediate strain rates) superplastic behaviour; region III (high strain rates) climb controlled dislocation creep. Although experimental results for region I are contradictory, there is agreement about the creep behaviour in regions II and III, which is usually expressed by double power law equation: snII QII QIII nIII e_ AII p exp AIII s exp (6) d kT kT
The specimen (Fig. 1) consisted of two silicon chips, which were connected by four ¯ip chip joints (each on one corner) to each other. SnAg and SnAgCu ¯ip chip bumps were printed 200 mm 200 mm Cu pads. Sn63Pb37 ¯ip chip bumps were deposited by electroplating on 100 mm 100 mm Cu pads. The joint shape was chosen hyperbolic. In contrast to the usual barrel shaped joints, in hyperbolic shaped joints, all strains and stresses will concentrate in the centre away from the interface to the under bump metalisation. That way, any interference with the interfacial region to the under bump metalisation is avoided. One of the problems that comes along which such type of real ¯ip chip specimen is to design a matching test setup for mechanical experiments. The imposed loading conditions should be reversal shear under isothermal conditions with appropriate cyclic strain rates. This requires a well controlled movement of the two silicon chips of the specimen against each other in such a way that all bending moments and out of plane forces are eliminated or at least minimised. This is realised through a symmetric grip con®guration of two identical specimens (Fig. 2). The test setup is driven by a piezoelectric translator, which provides a smooth movement with sub-nanometer resolution for a wide range of velocities. A force sensor is mounted between the inner grip and the
where d is the grain size, p the grain size exponent and AII, AIII the pre-factors. In region II, the SnPb eutectic has usually a low stress dependence with nII 2, when it exhibits superplastic deformation. 2.2. Creep in precipitation strengthened alloys The creep strength of two-phase or multiphase alloys can be dramatically higher then this of single phase alloys because of interactions between dislocations and metallurgical obstacles. Precipitated particles act as the most effective one. Theoretical models of dislocation movement in particle strengthened metals assume, that at high temperatures dislocations can undergo non-planar motion by climb, which allows the dislocation segment arrested at a particle bulge out of the slip plane and ®nally surmount the particle [3]. Since the traditional models of dislocation movement in metals were unable to explain high values of stress exponent n (upto 40) or unrealistic values of the apparent activation energy (Q), which were obtained from creep experiments on precipitation strengthened alloys, a creep strength increment sth s sm (s: applied stress; sm: creep strength of the referring particle-free metal matrix) was introduced to theory. With the empirical value of sth, the steady-state creep rate dependence on stress in precipitation strengthened alloys can be expressed as s s n th e_ A4 (7) E where n is the stress exponent for steady-state deformation of the reference material having the same composition as the matrix that the particles are dispersed in. The creep strength increment sth is no true threshold because it decreases with
Fig. 1. Flip chip specimens.
190
S. Wiese et al. / Sensors and Actuators A 99 (2002) 188±193
Fig. 2. Test setup.
piezoelectric translator. The translation direction is in line with the gravity axis, in order to avoid any bearing. Bearings cause frictional forces of at least 150 mN, which will not allow a high resolution force measurement. A high resolution displacement measurement is achieved by the application of a dual beam laser interferometer. Each laser beam is re¯ected at one of two neighbouring edges of the specimen silicon chips. In order to minimise errors due to temperature changes or vibrations, the laser interferometer heads are borne symmetrically to the specimen. The achieved true precision of the microtester is 20 nm for displacement and 2 mN for force. A more detailed description of the test setup and the experimental programme is given in [5]. 4. Results The steady-state creep data of eutectic SnAg/SnAgCu solders is plotted versus shear stress in Fig. 3. All data was experimentally determined at temperatures between 5 and 50 8C. Fig. 4 compares the corresponding microstructure of the ¯ip chip solder (V 1 10 12 m3) and compares it with
Fig. 3. Creep behaviour of Sn96.5Ag3.5/Sn95.5Ag4Cu0.5 solders.
that of an Sn95.5Ag4Cu0.5 bulk specimen (V 1 10 8 m3). The volume ratio is 1:10,000. The microstructure is composed of Ag3Sn intermetallics (dark) distributed in a b-Sn matrix (light). Intermetallics of Z-Cu6Sn5 appear also light and therefore, cannot been distinguished from the b-Sn matrix. Both specimens were soldered using SMT similar re¯ow pro®le. Although the differences in microstructure between both specimens enormous, it is very dif®cult to say, if these differences derive from volume effects or from differences in solder composition. Since the alloying content of Ag (4%) and Cu (0.5%) is relatively little, the solution of UBM metals (Au, Ni, Cu, etc.) during bumping and soldering processes may change the composition of the solder and thus the solidi®cation process changes signi®cantly, because the molten solder has moved away from the eutectic in the complex ternary SnAgCu system. The steady-state displacement rate data, that was gained on ¯ip chip specimens of eutectic SnPb solder at temperatures of 5 and 50 8C is plotted versus shear force in Fig. 5. All odd numbers correspond to specimens with an as cast microstructure (grain size <1 mm) and all even numbers correspond to specimens with a coarsened microstructure (grain size 10 mm). Force and displacement rate values were not converted into stress and strain rate values because the material model calibration was performed through FEM simulation [5]. The microstructure of Sn63Pb37 in a bulk solder volume (V 1 10 8 m3) and in a ¯ip chip joint (V 1 10 12 m3) is shown as a back scattered electron images in Fig. 6. Both specimens were soldered with a SMT similar re¯ow pro®le. What is common on both microstructures are lead-rich (light) clusters of globular shape with a diameter (d) between 5 and 20 mm. The rest of microstructure appears different. While the ¯ip chip cross-section shows a very homogeneous distribution of globular shaped lead-rich phases (light) in a tin-rich (dark) matrix (mean phase intercept approximately 0.9 mm), the cross-section of the bulk solder contains a non-uniform distribution of phases. The mean phase intercept starts at approximately 4 mm at coarse structured regions and ends at approximately 0.5 mm at ®ne
S. Wiese et al. / Sensors and Actuators A 99 (2002) 188±193
191
Fig. 4. Optical micrograph of Sn95.5Ag4Cu0.5 solder in a flip chip and a bulk specimen.
undercooling (DT 55 K) to nucleate, the nucleation of lead in the presence of tin starts with only little undercooling (DT 0:25 K) [6]. Since the undercooling that is needed for the homogeneous nucleation of lead on primary lead-rich crystals is very similar to the undercooling required for the heterogeneous nucleation of lead on primary lead-rich crystals, it does not matter if the composition of the solder is slightly hyper- or hypoeutectic, when the temperature falls below the eutectic point lead-rich phases start to grow until the composition of the liquid phase has been shifted to 80 wt.% Sn±20 wt.% Pb. At this composition the temperatures at which tin and lead-rich phase nucleates are very nearly the same [7]. Fig. 5. Creep behaviour of Sn63Pb37 solder.
structured regions. This very speci®c microstructural appearance of SnPb solder can be explained by the nonreciprocal nucleation behaviour of the bimaterial tin±lead system. While tin in the presence of lead needs a high
5. Discussion In order to compare the steady-state creep behaviour of Sn63Pb37, Sn96.5Ag3.5 and Sn95.5Ag4Cu0.5, the two governing parameters n and Q of Eqs. (1)±(7) should be
Fig. 6. Back scattered electron images of the Sn63Pb37 solder in a bulk and flip chip specimen.
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S. Wiese et al. / Sensors and Actuators A 99 (2002) 188±193 Table 2 Parameters for creep model (Eq. (9)) C1 (s 1) Sn96.5Ag3.5 Sn95.5Ag4Cu0.5
5 10 2 10
6 21
C2
C3 (MPa 1)
C4 (kJ/mol)
11 18
1 1
79.8 83.1
behaviour of precipitation strengthened materials is modelled by a simple power law C4 (9) e_ C1 C3 sC2 exp kT
Fig. 7. Comparison of creep of Sn96.5Ag3.5, Sn95.5Ag4Cu0.5 and Sn63Pb37 (all data points are corrected for T 300 K)
considered separately. For this reason the steady-state creep rates from the tests conducted at temperatures between 5 and 50 8C on ¯ip chip joints, were corrected for a temperature of T 27 8C using Eq. (5) and assuming a apparent activation energy. Fig. 7 shows the comparison of selected creep data of the three solders investigated. At room temperature Sn63Pb37 and Sn96.5Ag3.5 solder show nearly the same absolute creep rates at high stresses. At low stresses, however, Sn63Pb37 solder reduces its creep rate dependence on stress to an exponent of approximately n 2, whereas, a constant stress exponent of n 11 was observed for Sn96.5Ag3.5 solder. The low stress exponent of n 2 for Sn63Pb37 indicates that the solder in ¯ip chip joints deforms at least partially superplastically, when stresses are low. Modelling of creep behaviour of (near) eutectic tin lead solder (Table 1) was performed using the sinh formulation from Eq. (4) C4 C2 e_ C1 sinh C3 s exp (8) kT The high stress exponent of n 11 for Sn96.5Ag3.5 can be explained by a high value of sth (Eq. (7)). Since the cooling rate during solidi®cation is very high for a ¯ip chip joint, it is very likely that the solder in the ¯ip chip joint contains a high number of very small Ag3Sn precipitates. In contrast, the same solder in a bulk specimen that was solidi®ed at low cooling rate contains a smaller number of larger Ag3Sn precipitates (Fig. 4). At the same volume fraction of Ag3Sn a higher number of smaller particles acts as a more potent obstacle to the movement of dislocations than a smaller number of larger particles, which results in high values of for the solder in ¯ip chip joints. If creep Table 1 Parameters for creep model (Eq. (8))
Sn63Pb37
C1 (s 1)
C2
C3 (MPa 1)
C4 (kJ/mol)
10
2
0.2
44.9
the derivation of Eq. (7) gives ns @sth 1 C 2 n0 s sth @s T
(10)
Since n is a constant and is given by the matrix metal (stress exponent of b-Sn n 4 5), it can be concluded from Eq. (10), that n0 increases with increasing sth. Therefore, it can be assumed that the creep properties of precipitation strengthened material like Sn96.5Ag3.5 show a strong dependence on the cooling conditions that in¯uence the microstructure of the solder. Since it is very dif®cult to determine the value of sth experimentally it was preferred to model the creep behaviour of Sn96.5Ag3.5 solder (Table 2) with the simple power law from Eq. (9). Sn95.5Ag4Cu0.5 solder in ¯ip chip joints showed a stress exponent of n 18 (actually n0 18 for Eq. (10)) which is signi®cantly higher than that found on the Sn96.5Ag3.5 material. In addition, the ternary eutectic of the SnAgCu alloy showed much lower absolute creep rates than the binary eutectic of the SnAg material. This extreme increase in creep resistance may result from tiny (5 and 50 nm) precipitates of Z-Cu6Sn5 in the b-tin matrix [8,9]. Such small particles act as potent obstacles for dislocation movement and thus the creep resistance of an SnAgCu solder can be increased dramatically. For creep modelling of the Sn95.5Ag4Cu0.5 (Table 2) ¯ip chip solder the simple power law (Eq. (9)) as for the Sn96.5Ag3.5 was chosen. The apparent activation energy of creep Q (Eq. (5)) found on ¯ip chip specimens of Sn63Pb37 solder is in accordance with what is reported in literature. However, there are strong differences between literature values [10±14] of activation energy Q and these found in this study for Sn96.5Ag3.5 and Sn95.5Ag4Cu0.5 solder. These differences result supposedly from the fact that the ®t function (Eq. (9)) include implicit the creep resistance increment. Therefore, the parameter C4 (apparent activation energy) in the Eq. (9) corresponds to a corrected activation energy Q0 , that results from Eqs. (5) and (7) [4] n 1 dE n @sth Q0 Q kT 2 (11) E dT s sth @s s where Q is the apparent activation energy for the creep mechanisms in the precipitation strengthened material,
S. Wiese et al. / Sensors and Actuators A 99 (2002) 188±193
which is corrected by the temperature dependence of Young's modulus and sth. It is reasonable to assume that the temperature dependence of sth is similar to that of E. Since dE=dT < 0, Q0 increases when the difference s sth becomes smaller. This means larger values of n0 (C2), which result from higher correspond to larger values of Q0 (C4). 6. Conclusions The steady-state creep rate was determined by reversible constant load shear tests on Sn96.5Ag3.5, Sn95.5Ag4Cu0.5 and Sn63Pb37, ¯ip chip solder joints. The stress exponent found on Sn96.5Ag3.5 solder (n 11) and Sn95.5Ag4Cu0.5 solder (n 18) was signi®cantly higher than that found on Sn63Pb37 solder (n 2) in the low stress region. Also the apparent activation energy was higher for the Sn96.5Ag3.5 (Q 79:8 kJ/mol) and the Sn95.5Ag4Cu0.5 (Q 83:1 kJ/ mol) solder then for the Sn63Pb37 (Q 44:9 kJ/mol) solder. Microstructural analysis and theoretical assumptions turn out that the high values for n and Q that were found on the Sn96.5Ag3.5 and Sn95.5Ag4Cu0.5 solder might be caused by small precipitates of Ag3Sn and Z-Cu6Sn5 intermetallics that are ®nely dispersed in the b-tin matrix. References [1] C.E. Pearson, The viscous properties of extruded eutectic alloys of lead±tin and bismuth±tin, J. Inst. Metals 54 (1934) 111±124. [2] D.H. Avery, W.A. Backofen, Trans. Am. Soc. Metals 58 (1965) 551.
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[3] B. Reppich, Particle strengthening, in: R.W. Cahn, P. Haasen, E.J. Kramer (Eds.), Materials Science and TechnologyÐA Comprehensive Treatment, Vol. 6, VCH, Weinheim, 1993, pp. 311±355. [4] J. Cadek, Creep in Metallic Materials, Elsevier, Amsterdam, 1988. [5] S. Wiese, F. Feustel, S. Rzepka, E. Meusel, Experimental characterisation of material properties of Sn63Pb37 flip chip solder joints, in: D.J. Belton, M. Gaynes, E.G. Jacobs, R. Pearson, T. Wu (Eds.), Proceedings of the Tenth MRS Symposium on Electronic Packaging Materials Science, Vol. 515, Materials Research Society, San Francisco, 14±16 April 1998, pp. 233±238. [6] B.E. Sundquist, L.F. Mondolfo, Heterogeneous nucleation in the liquid-to-solid transformation in alloys, Trans. Metallurg. Soc. AIME 221 (1961) 157±163. [7] J.H. Hollomon, D. Turnbull, Solidification of lead±tin alloy droplets, J. Metal Trans. AIME 3 (1951) 803±805. [8] L.E. Felton, K. Rajan, P.J. Ficalora, P. Singh, n-Cu6Sn5 precipitates in Cu/PbSn solder joints, Scripta Metallurgica et Materialia 25 (1991) 2329±2333. [9] L. Xiao, J. Liu, Z. Lai, L. Ye. A. ThoÈlen, Characterization of mechanical properties of bulk lead-free solders, in: Proceedings of the International Symposium on Advanced Packaging Materials, Braselton, GA, 6±8 March 2000, pp. 145±151. [10] E.I. Stromswold, Characterization of eutectic tin±silver solder joints, Dissertation, University of Rochester, 1993. [11] R. Darveaux, K. Banerji, Constitutive relations for tin-based solder joints. IEEE Transactions on CHMT 15 (1992) 1013±1024. [12] Z. Guo, Y.H. Pao, H. Conrad, Plastic deformation kinetics of Sn95.5Cu4Ag0.5 solder joints, J. Electron. Packaging 117 (1995) 100±104. [13] D.R. Frear, The mechanical behaviour of interconnect materials for electronic packaging, J. Metals 48 (1996) 49±53. [14] A. Schubert, H. Walter, A. Gollhardt, B. Michel, Materials mechanics and reliability issues of lead-free solder interconnects, in: Proceedings of the Second Conference on Benefiting from Thermal and Mechanical Simulation in Micro-ElectronicsÐEuroSimE 2001, Paris, 9±11 April 2001, pp. 171±178.