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Tensile fracture of tin–lead solder joints in copper
Materials Science and Engineering A 379 (2004) 277–285
Tensile fracture of tin–lead solder joints in copper K.H. Prakash, T. Sritharan∗ School of Materials Engineering, Nanyang Technological University, Nanyang Avenue, Singapore-639798, Singapore Received 4 September 2003; received in revised form 10 February 2004
Abstract Tensile testing was carried out at constant cross-head speed of 0.5 mm/min on specimens of two Cu blocks of length 45 mm with rectangular area of cross-section (10 mm × 3 mm) soldered using Sn–Pb solder alloys with 200–300 m thick. Four compositions of Sn–Pb alloys were used. The testing specimens were aged at 175 ◦ C in an oven with a temperature measuring accuracy of ±1 ◦ C for different times of up to 16 days to study the effect of isothermal ageing on the strength of the joints. The fracture surface was analysed both in the plan and cross-sectional views to identify the fracture location in the solder joint. It was identified that the morphology of the intermetallic compound (IMC) layer plays a role on the strength of the joint, especially when it fails through the solder/IMC interface. Also, the Pb-content in the solder alloy has a profound effect on the fracture path. © 2004 Elsevier B.V. All rights reserved. Keywords: Solder joint reliability; Interface intermetallics; Tensile fracture
1. Introduction The solder joints connect the electronic components to the board electrically as well as mechanically. In surface mount technology (SMT) for instance, the components get mechanical support solely from the solder joints. Failure of solder joints could occur by mechanical shock during manufacture, transportation, etc., high-cycle fatigue (vibration during in-service condition) or thermomechanical fatigue (environmental temperature changes, change in temperature during power on/off cycles, etc.). Of these, the life under thermomechanical fatigue is an important reliability indicator. Extensive work has been done on the isothermal fatigue testing and the thermomechanical fatigue testing of solder joints [1–6]. Shear testing is done to determine the mechanical integrity of the joint, which is an indirect indicator of the electrical connection as well. A substantial amount of work is available in the literature on shear testing of solder joints using shear lap joints [7–10]. Studies on simultaneous measurement of electrical resistance and mechanical deformation of solder joints show that the conductivity was not affected when the joint reached its yield strength but decreased abruptly after complete debonding [11]. ∗
Corresponding author. Tel.: +65-6790-4586; fax: +65-6790-9081. E-mail address: [email protected] (T. Sritharan).
0921-5093/$ – see front matter © 2004 Elsevier B.V. All rights reserved. doi:10.1016/j.msea.2004.02.049
Shear testing is a common quality control test, particularly in ball grid arrays. This test is easy and quick to use on real components but is not valuable when one is investigating the failure mechanism. The results are difficult to interpret in terms of fundamental strength properties because the prevailing stress conditions are not known accurately, and the friction between the shearing surfaces also contributes to the failure load. Further, the frictional damage impairs fracture surface investigations. Thus, lap shear and tensile tests are preferred by researchers. These have to be done on large, specially made experimental joints but could give reliable results regarding their strength. In addition tensile fracture surfaces are better preserved and could lend more evidence towards the failure mechanism. In tensile tests the stress is more uniform across the section and is the same the on all the constituents namely the solder, the substrate and the interfaces. As a result the failure should occur at the weakest constituent of the joint. Hence, tensile testing is used to assess the joint strength in this study.
2. Experimental procedure Cu samples of length 45 mm were cut from a commercial grade, extruded Cu bars with rectangular cross-section (10 mm × 3 mm). These were etched in 50% nitric acid to
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Table 1 Solder compositions, their freezing ranges and the soldering temperatures used Solder composition (wt.%)
Solidus temperature (◦ C)
Liquidus temperature (◦ C)
Soldering temperature (◦ C)
100Sn 82Sn18Pb 63Sn37Pb 27Sn73Pb
232 183 183 183
232 200 183 250
242 210 193 260
remove the surface oxide layer. The mating surfaces were fluxed immediately with commercial Rosin mildly activated (RMA) flux and the rest of the surfaces were wrapped with Al foil to prevent them being wetted by molten solder. Solder alloys were rolled into thin sheets of thickness 1 mm and sliced into pieces that approximately covered the mating surface area of the Cu samples. Then, the sliced solder pieces were placed between the mating surfaces of two Cu samples in an Al mould and were heated in a furnace to a temperature 10 ± 2 ◦ C above the liquidus of the solder. The solder compositions and the soldering temperatures used are given in Table 1. After holding in the molten state for 5 min, the samples were gently pushed towards each other to obtain a good joint and were cooled in the furnace. It was found that the tensile testing specimens thus prepared resulted in joints with solder of thickness of between 200 and 300 m, as schematically shown in Fig. 1. The specimens were annealed at 175 ◦ C in an oven with a temperature measuring accuracy of ±1 ◦ C for 1, 3, 7 and 16 days. For each condition, four samples were made; one was used to study the interface intermetallic compound (IMC) and the other three were used for tensile testing. Tensile testing was done in an INSTRON 4206 tensile tester at room temperature with a constant cross-head speed of 0.5 mm/min. The untested sample was prepared in the standard manner for cross-section metallographic investigation and the IMC layer thickness was measured using an image analysing software on the digitised scanning electron micrographs. For each sample measurements were made in two regions that are far apart. In each of these regions, 30 measurements were made at regular intervals of 5 mm. The averages of the 60 measurements are used in the analysis. The IMC layer thickness at each measurement position was taken as the linear distance between the Cu substrate and the top of the IMC layer at that locality.
Fig. 1. Dimensions of the tensile testing specimen used in this study. The dimensions are in millimetres and the hatched region represents the solder alloy.
In fractured samples, the fracture surfaces were protected by application of a thick layer of metallic glue and were cold mounted and polished as usual. A piece of Cu was attached to the glue to provide stability to the edges during polishing.
3. Results 3.1. Morphology of the IMC layer The cross-sections of the samples in the as-soldered condition showed only a single layer of -phase (Cu6 Sn5 ) at the interface. Fig. 2 shows micrographs of cross-sections of the joint. The morphology of the -phase layer depended on the Pb-content of the solder. 100Sn solder gave a cellular morphology and it gradually transformed to scallop type as the Pb-content of the solder increased. The morphologies and growth kinetics of the IMC layers when the solders are in the molten state had been analyzed and published previously by the authors [12]. Upon annealing at 175 ◦ C, the ε-phase (Cu3 Sn) layer appeared between -phase and Cu. The -phase/solder interface became relatively planar with increase in annealing time except in 100Sn. Fig. 3 shows the cross-sections after annealing at 175 ◦ C for 16 days. It is clear that the -phase/solder interface becomes more planar as the Pb-content of the solder increases. The morphologies of the IMC layers and their growth rates during annealing in the solid state have been published elsewhere [13]. 3.2. Solder joint strength Fig. 4 shows the load versus displacement curves for the solder joints in the as-soldered condition tested at room temperature. Because the thickness of solder in the joint was not strictly constant in all samples, stress–strain curves are not compared. Note that the curves in Fig. 4 are almost linear to failure, indicating the failure of the joints mostly occurred within the elastic strains or with very little plastic strains. Annealed joints also showed similar behaviour. Any plastic deformation must have been restricted to the thin solder between the two Cu blocks. The values of fracture strength (average of maximum stress) obtained for different solders are given in Table 2 for each annealing time, along with their corresponding total IMC layer ( + ε) thickness. A graphical plot of fracture strength against annealing time is shown in Fig. 5. Error bars are significantly large to clearly identify any trends in the graphs. Fracture strengths of Cu/100Sn joints show the widest scatter and no trend could be observed. Strengths of Cu/82Sn18Pb joints show less scatter compared to Cu/100Sn joints. Those of other two solders appear to mildly decrease with annealing time. The large scatter in fracture strength of Cu/100Sn joints could be attributed to the morphology of the IMC layer. The undulating, non-planar -phase/solder interface could have led to the large scatter in the strength of the 100Sn solder joints if the fracture had propagated along
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Fig. 2. Sectional view showing the morphology of the -phase layer at the interface in the as-soldered joints: (a) 100Sn; (b) 82Sn18Pb; (c) 63Sn37Pb and (d) 27Sn73Pb. Labelling is done only in (a) to show the location of the solder, the Cu and the intermetallic. Orientations of the samples in (b–d) are the same as in (a).
this interface. In other solders, the IMC morphology is more planar, as evident from Fig. 3, resulting in less scatter. Several workers have reported that the increase in IMC layer thickness decreases the joint strength [14,15]. For example, Quan et al. [14] reported a 50% decrease in tensile strength of Cu/63Sn37Pb and Cu/5Sn95Pb joints when the total IMC layer thickness increased from an initial 4 m to final thickness of 15 and 11 m, respectively. Our results show a lower decrease in strength of 18.4 and 22% with increase in IMC layer thickness in Cu/63Sn37Pb and Cu/27Sn73Pb joints, respectively. Fig. 6 shows a plot of fracture strength against the Pb-content of the solder at different annealing times. It is evident that the joint strength reaches a maximum at the eutectic composition of 63Sn37Pb. The 100Sn solder has the lowest strength. It is known that the tensile strength of the bulk Sn–Pb solders increases with Pb-content until the eutectic composition (63Sn37Pb), and thereafter decreases with further increase in Pb-content. For example, the tensile strength of 5Sn95Pb is reported to be about 29 MPa, that of 63Sn37Pb is 46 MPa and that of 100Sn is 12 MPa at room temperature [16]. Our results show a similar trend in the
variation of joint strength with Pb-content of the solder. The authors in Ref. [16] did not mention the strain rate at which they performed the tensile testing. Hence we could only compare trends in the effect of Pb-content on the strength, not the actual strengths of the joints. Moreover, there are more potential failure paths in the solder joints than the bulk solders which can only fail through obviously the bulk solder. 3.3. Fracture surface analysis In the uniaxial tension, solder joints must fail at the weakest plane in the three constituents namely, the solder, the IMC or the Cu bar, or at one of the interfaces between these constituents. An examination of the fracture path will thus give an insight to the relative strengths of the constituents and interfaces. As the strength of Cu is considerably larger than those of the others, none of the samples failed through Cu. Fracture path was observed to go through various constituents and locations depending on the annealing time, solder composition and the nature of the interfaces. These will be discussed below.
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Fig. 3. Sectional view showing -phase and ε-phase layers at the interface in joints annealed at 175 ◦ C for 16 days: (a) 100Sn; (b) 82Sn18Pb; (c) 63Sn37Pb and (d) 27Sn73Pb. Labelling is done only in (a) to show the location of the solder, the Cu and the intermetallics. Orientations of the samples in (b–d) are the same as in (a).
3.3.1. The high-Pb solder 27Sn73Pb Fracture was found to occur totally through the bulk solder, in a ductile manner in the high-Pb solder 27Sn73Pb under some processing conditions. Fig. 7 shows the fracture surface and a sectional view of a Cu/27Sn73Pb joint in the as-soldered condition. The feature of microvoid coalescence mechanism typical of ductile failure is apparent in Fig. 7a, while the sectional view of Fig. 7b shows that the fracture path is through the bulk solder. After annealing for 1 day
at 175 ◦ C, the fracture looked similar to that in Fig. 7. The samples annealed for 16 days at 175 ◦ C, however, showed a faceted appearance with no sign of ductility, indicating a different fracture mechanism and path. An example of the fracture surface at this condition is shown in Fig. 8. 3.3.2. The pure Sn solder 100Sn The 100Sn solder is also a weak material. Then, a ductile failure through the solder could be expected in the
Table 2 Fracture strengths and total IMC layer thicknesses for each solder composition and annealing time Time (days)
0 1 3 7 16
Pure Sn
82Sn18Pb
63Sn37Pb
27Sn73Pb
Total IMC layer thickness (m)
Maximum fracture strength (MPa)
Total IMC layer thickness (m)
Maximum fracture strength (MPa)
Total IMC layer thickness (m)
Maximum fracture strength (MPa)
Total IMC layer thickness (m)
Maximum fracture strength (MPa)
5.37 7.63 8.27 9.93 11.81
34.9 44.6 30.0 60.2 34.8
4.22 6.59 8.42 10.08 12.24
65.7 77.1 76.6 66.4 73.4
2.47 4.60 6.13 8.27 11.29
85.8 84.0 85.8 78.8 70.0
2.81 5.08 6.12 7.55 10.38
80.1 65.8 63.9 60.8 62.5
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Fig. 4. Load vs. displacement curves of solder joints made with the four solder in the as-soldered condition. The curve for 82Sn18Pb solder is hidden by the curve for 63Sn37Pb solder, as arrowed in the figure.
Cu/100Sn joints too. Fig. 9 shows the fracture surfaces in Cu/100Sn joints in the as-soldered, 1 and 16 days annealed conditions. The as-soldered sample (Fig. 9a) and 1 day annealed sample (Fig. 9b) show typical ductile failure features in most parts of the fracture but some faceting is evident. The 16 days annealed sample (Fig. 9c) shows a faceted surface similar to that in Fig. 8 for Cu/27Sn73Pb joint annealed for 16 days. The faceted fracture dominant in Fig. 9c is evident at the base of the circular microvoid features in Fig. 9b. A sectional view of a fractured sample of the joint annealed for 1 day (Fig. 10a) clearly shows that, although the fracture path is near the -phase/solder interface, in
Fig. 5. Plot of fracture strength against the annealing time for the joints made of all four solders.
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Fig. 6. Plot of fracture strength against the Pb-content of the solder at various annealing durations.
order to remain perpendicular to the tensile stress, it has propagated through the solder in some areas and through the -phase in others. Thus, the faceted features evident at the base of the microvoids in Fig. 9b are probably the cleaved -phase crystals when the fracture propagates. When the fracture is through the solder, it fails by a ductile, microvoid coalescence mechanism. Even in the as-soldered joint, some -phase crystals were identified at the base of the microvoids. A sectional view of the 16 days annealed joint, shown in Fig. 10b, shows that the fracture path is predominantly through the -phase layer, but is near the -phase/ solder interface. Thus, in 100Sn solder joints, the fracture occurs in the vicinity of the -phase/solder interface in all conditions. The uneven -phase/solder interface typical in these alloys (see Fig. 2) may have nucleated the microvoids during the tensile test, encouraging failure to remain near the interface. However, if the fracture is to occur exactly on the interface, it will have to change direction abruptly and frequently to propagate as the interface is uneven, which would be energetically unfavourable. Fracture through the -phase crystals appears to have been by cleavage, giving the faceted appearance. This is possible since it was previously shown that the -phase grains in the IMC layers have a preferred crystallographic orientation [17]. Then cleavage planes in neighbouring -phase grains could be nearly parallel for the fracture to propagate easily through the IMC. The relative area fractions of the solder and IMC on the fracture surface will vary from sample to sample as the interface morphology would not be identical even if the external processing conditions are nominally the same. This may have caused the high scatter in the fracture strength in joints with 100Sn solder. When annealed, the IMC layer grows and becomes more planar reducing the scatter in the fracture strength.
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Fig. 7. Fracture in a Cu/27Sn73Pb joint in the as-soldered condition: (a) the fracture surface (b) a sectional view of one side. The fracture (arrowed) is through the bulk solder away from the interface.
3.3.3. The low-Pb solder 82Sn18Pb and eutectic solder 63Sn37Pb The fracture surfaces of the as-soldered joints with these solders are shown in Fig. 11. Some ductility is evident. Some -phase grains are visible on the fracture surfaces. The exposure of -phase grains could be because of -phase/solder
Fig. 9. Fracture surfaces in Cu/100Sn joints: (a) as-soldered; (b) annealed for 1 day at 175 ◦ C and (c) annealed for 16 days at 175 ◦ C.
Fig. 8. Fracture surface of a Cu/27Sn73Pb joint in annealed at 175 ◦ C for 16 days.
interface decohesion caused by microvoid formation and growth. Thus, the fracture has occurred in the vicinity of the -phase/solder interface. The fracture surfaces and the corresponding sectional views of samples annealed for 1 and 16 days are shown in Fig. 12. It is clear that the fracture had occurred through the IMC layer exposing the cleaved -phase grains. The sectional view shows that the fracture occurred through the -phase layer, but near the -phase/ε-phase interface. In the as-soldered condition, fracture occurred near the
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Fig. 10. Sectional view of a fracture surface in a Cu/100Sn joint annealed at 175 ◦ C for (a) 1 day and (b) 16 days. Fracture path is arrowed. Intermetallics appear grey. Cu is not visible in (a) as it is in the other half of the fractured sample. In (b) one Cu piece and its interface with solder are visible but the fracture is at the interface with the other Cu piece that is not in this half of the sample.
-phase/solder interface. When annealed for 16 days it appeared to have transferred to the -phase/ε-phase interface.
4. Discussion Isothermal annealing simultaneously coarsens the microstructure of the solder and increases the IMC layer thickness. Coarsening of microstructure is generally expected to decreases the tensile strength. However, according to a study by Takemoto et al. [18] in Sn–Pb alloys, the microstructure coarsening appears to have little effect on its tensile strength. For example, in eutectic 63Sn37Pb solder, they reported a tensile strength of 44.4 MPa in an as-cast sample which, when annealed at 120 ◦ C for 7 and 21 days, decreased slightly to 42.2 and 42.5 MPa, respectively. Thus, the tensile strength of the bulk solder appears to remain constant after a slight initial decrease. However, we cannot attribute the decrease in strength observed on annealing our joints to the solder microstructure coarsening as we have shown that a transition occurs in the fracture path and
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Fig. 11. Fracture surface of the (a) Cu/82Sn18Pb and (b) Cu/63Sn37Pb joints in the as-soldered condition.
mode. As evident from Table 2, the IMC layer grew to 10 m or higher in all joints when annealed at 175 ◦ C for 16 days. The fracture also occurred through the -phase layer in these samples with no features of ductility. Then, strengths obtained in these samples must reflect the strength of the IMC layer irrespective of the solder composition. This appears to be true since the strengths of all 16 days annealed samples appear to approach a similar value in the range 62–73 MPa, except in 100Sn (Fig. 5). The reason for the different behaviour of the Cu/100Sn could be the rough morphology of the -phase/solder interface. In comparison, the as-soldered samples of 100Sn, 82Sn18Pb and 63Sn37Pb solder joints generally fractured near the -phase/solder interface and the joint strength increased with increasing Pb-content (Fig. 6). For example, the fracture strength of Cu/63Sn37Pb (85.8 MPa) is higher than that of Cu/82Sn18Pb joint (65.7 MPa) although their respective fracture surfaces appear similar. A contributory factor for this difference could be the cohesive strength of the interface. It is reported that the surface tension of molten pure Pb (0.47 N/m K) is lower than that of pure Sn (0.55 N/m K) at their respective melting temperatures [19]. This implies that molten Pb will wet solids more easily than molten Sn. If we consider wetting as an indicator of the
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Fig. 12. Fracture in a Cu/82Sn18Pb joint annealed at 175 ◦ C for 1 day: (a) the fracture surface (b) a sectional view. Fracture in a Cu/63Sn37Pb joint annealed at 175 ◦ C for 16 days: (c) the fracture surface (d) a sectional view. Sectional views are labelled to show the fracture path, the solder and the intermetallics.
cohesive strength, then the cohesive strength of the -phase/Pb interface should be higher than that of the -phase/Sn interface. As the fraction of -phase/Pb interface increases with increasing Pb-content of the solder, the average cohesive strength of the -phase/solder interface could also increase. The high-Pb solders will then have a high interface cohesive strength and thus the fracture should occur away from the interface as observed in as-soldered Cu/27Sn73Pb joint. The Pb containing solder joints fractured through the -phase after annealing. Annealing thickens the IMC layer and promotes accumulation of Pb at the -phase/solder interface as Sn reacts with Cu to form the IMC. Cohesive strength of the interface could then increase with annealing. Therefore, a combination of thicker IMC layer and stronger -phase/solder interface would favour fracture through -phase grains instead of decohesion at the interface or failure through the bulk solder. 5. Conclusions Tensile testing of Cu/solder joints at room temperature and at constant cross-head speed were performed with four solder compositions 100Sn, 82Sn18Pb, 63Sn37Pb and
27Sn73Pb for annealing times up to 16 days at 175 ◦ C. The fracture strengths and the fracture paths were determined and are related to the microstructure. The following conclusions can be made from this study: 1. The as-soldered joints have a layer of -phase. The morphology and thickness of the -phase varied depending on the solder composition and annealing time. The morphology was uneven in 100Sn solder but became scalloped as the Pb-content increased. 2. The fracture strength of the joint, in general, increased with increasing Pb-content in the solder alloy up to the eutectic composition (63Sn37Pb) and then decreased with further increase in Pb. 3. The fracture strength of Cu/100Sn joints showed a large scatter with annealing time compared to those of the other solders. This could be attributed to the uneven interface between the -phase and the solder in these joints. 4. In the as-soldered condition, fracture occurred near the -phase/solder interface, except in Cu/27Sn73Pb joints. The fracture mechanism in these joints appears to be partly by microvoid coalescence in the bulk solder, partly by cleavage of -phase grains and partly by decohesion at the interface. In Cu/27Sn73Pb joints, however, fracture
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occurred totally by microvoid coalescence through the bulk solder in the as-soldered condition. 5. Annealing appears to transfer the fracture path to the -phase layer. The 16 days annealed samples predominantly failed by cleavage of the -phase grains. The fracture strengths of these samples also appear to converge to a value in a small band irrespective of the solder composition. 6. Increasing Pb-content in the solder could improve the cohesive strength of the -phase/solder interface and contribute to the increase in fracture strength of the joint.
References [1] [2] [3] [4]
D. Yao, J.K. Shang, IEEE Trans. CPMT-B 19 (1996) 154. H.D. Solomon, IEEE Trans. CHMT 9 (1986) 423. X. Zhang, S.-W. Ricky Lee, Y.H. Pao, Trans. ASME 122 (2000) 200. H. Akay, H. Zhang, N. Paydar, in: Advances in Electronic Packaging, Proceedings of the Pacific Rim/ASME International Intersociety Electronic and Photonic Packaging Conference, INTERpack’97, vol. 2, 1997, p. 1583.
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[5] W.W. Lee, L.T. Nguyen, G.S. Selvaduray, Microelectron. Reliab. 40 (2000) 231. [6] X. Ma, Y. Qian, F. Yoshida, J. Alloys Compd. 334 (2002) 224. [7] Z. Guo, A.F. Sprecher Jr., H. Conrad, M. Kim, in: Proceedings of the Conference on Materials Developments in Microelectronic Packaging: Performance and Reliability, Montreal, PQ, ASM International, August 1991, p. 155. [8] W.J. Tomlinson, A. Fullylove, J. Mater. Sci. 27 (1992) 5777. [9] R.W. Wild, in: Proceedings of the National Electronic Packaging Conference, Anaheim, CA, Industrial and Scientific Conference Management, Chicago, IL, 1974, p. 105. [10] M. Warwick, Brazing Soldering 8 (Spring) (1985) 20. [11] X. Luo, D.D.L. Chung, J. Mater. Sci. 34 (1999) 273. [12] K.H. Prakash, T. Sritharan, Acta Mater. 49 (2001) 2481. [13] K.H. Prakash, T. Sritharan, J. Electron. Mater. 32 (2003) 939. [14] L. Quan, D.R. Frear, D. Grivas, J.W. Morris Jr., J. Electron. Mater. 16 (1987) 203. [15] B.S. Chiou, J.H. Chang, J.G. Duh, IEEE Trans. CPMT-B 18 (1995) 537. [16] Website: http://www.boulder.nist.gov/div853/lead%20free/part1.html. [17] K.H. Prakash, T. Sritharan, J. Electron. Mater. 31 (2002) 1250. [18] T. Takemoto, A. Matsunawa, M. Takahashi, J. Mater. Sci. 32 (1997) 4077. [19] J. Glazer, Int. Mater. Rev. 40 (1995) 65.
Tensile fracture of tin–lead solder joints in copper K.H. Prakash, T. Sritharan∗ School of Materials Engineering, Nanyang Technological University, Nanyang Avenue, Singapore-639798, Singapore Received 4 September 2003; received in revised form 10 February 2004
Abstract Tensile testing was carried out at constant cross-head speed of 0.5 mm/min on specimens of two Cu blocks of length 45 mm with rectangular area of cross-section (10 mm × 3 mm) soldered using Sn–Pb solder alloys with 200–300 m thick. Four compositions of Sn–Pb alloys were used. The testing specimens were aged at 175 ◦ C in an oven with a temperature measuring accuracy of ±1 ◦ C for different times of up to 16 days to study the effect of isothermal ageing on the strength of the joints. The fracture surface was analysed both in the plan and cross-sectional views to identify the fracture location in the solder joint. It was identified that the morphology of the intermetallic compound (IMC) layer plays a role on the strength of the joint, especially when it fails through the solder/IMC interface. Also, the Pb-content in the solder alloy has a profound effect on the fracture path. © 2004 Elsevier B.V. All rights reserved. Keywords: Solder joint reliability; Interface intermetallics; Tensile fracture
1. Introduction The solder joints connect the electronic components to the board electrically as well as mechanically. In surface mount technology (SMT) for instance, the components get mechanical support solely from the solder joints. Failure of solder joints could occur by mechanical shock during manufacture, transportation, etc., high-cycle fatigue (vibration during in-service condition) or thermomechanical fatigue (environmental temperature changes, change in temperature during power on/off cycles, etc.). Of these, the life under thermomechanical fatigue is an important reliability indicator. Extensive work has been done on the isothermal fatigue testing and the thermomechanical fatigue testing of solder joints [1–6]. Shear testing is done to determine the mechanical integrity of the joint, which is an indirect indicator of the electrical connection as well. A substantial amount of work is available in the literature on shear testing of solder joints using shear lap joints [7–10]. Studies on simultaneous measurement of electrical resistance and mechanical deformation of solder joints show that the conductivity was not affected when the joint reached its yield strength but decreased abruptly after complete debonding [11]. ∗
Corresponding author. Tel.: +65-6790-4586; fax: +65-6790-9081. E-mail address: [email protected] (T. Sritharan).
0921-5093/$ – see front matter © 2004 Elsevier B.V. All rights reserved. doi:10.1016/j.msea.2004.02.049
Shear testing is a common quality control test, particularly in ball grid arrays. This test is easy and quick to use on real components but is not valuable when one is investigating the failure mechanism. The results are difficult to interpret in terms of fundamental strength properties because the prevailing stress conditions are not known accurately, and the friction between the shearing surfaces also contributes to the failure load. Further, the frictional damage impairs fracture surface investigations. Thus, lap shear and tensile tests are preferred by researchers. These have to be done on large, specially made experimental joints but could give reliable results regarding their strength. In addition tensile fracture surfaces are better preserved and could lend more evidence towards the failure mechanism. In tensile tests the stress is more uniform across the section and is the same the on all the constituents namely the solder, the substrate and the interfaces. As a result the failure should occur at the weakest constituent of the joint. Hence, tensile testing is used to assess the joint strength in this study.
2. Experimental procedure Cu samples of length 45 mm were cut from a commercial grade, extruded Cu bars with rectangular cross-section (10 mm × 3 mm). These were etched in 50% nitric acid to
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Table 1 Solder compositions, their freezing ranges and the soldering temperatures used Solder composition (wt.%)
Solidus temperature (◦ C)
Liquidus temperature (◦ C)
Soldering temperature (◦ C)
100Sn 82Sn18Pb 63Sn37Pb 27Sn73Pb
232 183 183 183
232 200 183 250
242 210 193 260
remove the surface oxide layer. The mating surfaces were fluxed immediately with commercial Rosin mildly activated (RMA) flux and the rest of the surfaces were wrapped with Al foil to prevent them being wetted by molten solder. Solder alloys were rolled into thin sheets of thickness 1 mm and sliced into pieces that approximately covered the mating surface area of the Cu samples. Then, the sliced solder pieces were placed between the mating surfaces of two Cu samples in an Al mould and were heated in a furnace to a temperature 10 ± 2 ◦ C above the liquidus of the solder. The solder compositions and the soldering temperatures used are given in Table 1. After holding in the molten state for 5 min, the samples were gently pushed towards each other to obtain a good joint and were cooled in the furnace. It was found that the tensile testing specimens thus prepared resulted in joints with solder of thickness of between 200 and 300 m, as schematically shown in Fig. 1. The specimens were annealed at 175 ◦ C in an oven with a temperature measuring accuracy of ±1 ◦ C for 1, 3, 7 and 16 days. For each condition, four samples were made; one was used to study the interface intermetallic compound (IMC) and the other three were used for tensile testing. Tensile testing was done in an INSTRON 4206 tensile tester at room temperature with a constant cross-head speed of 0.5 mm/min. The untested sample was prepared in the standard manner for cross-section metallographic investigation and the IMC layer thickness was measured using an image analysing software on the digitised scanning electron micrographs. For each sample measurements were made in two regions that are far apart. In each of these regions, 30 measurements were made at regular intervals of 5 mm. The averages of the 60 measurements are used in the analysis. The IMC layer thickness at each measurement position was taken as the linear distance between the Cu substrate and the top of the IMC layer at that locality.
Fig. 1. Dimensions of the tensile testing specimen used in this study. The dimensions are in millimetres and the hatched region represents the solder alloy.
In fractured samples, the fracture surfaces were protected by application of a thick layer of metallic glue and were cold mounted and polished as usual. A piece of Cu was attached to the glue to provide stability to the edges during polishing.
3. Results 3.1. Morphology of the IMC layer The cross-sections of the samples in the as-soldered condition showed only a single layer of -phase (Cu6 Sn5 ) at the interface. Fig. 2 shows micrographs of cross-sections of the joint. The morphology of the -phase layer depended on the Pb-content of the solder. 100Sn solder gave a cellular morphology and it gradually transformed to scallop type as the Pb-content of the solder increased. The morphologies and growth kinetics of the IMC layers when the solders are in the molten state had been analyzed and published previously by the authors [12]. Upon annealing at 175 ◦ C, the ε-phase (Cu3 Sn) layer appeared between -phase and Cu. The -phase/solder interface became relatively planar with increase in annealing time except in 100Sn. Fig. 3 shows the cross-sections after annealing at 175 ◦ C for 16 days. It is clear that the -phase/solder interface becomes more planar as the Pb-content of the solder increases. The morphologies of the IMC layers and their growth rates during annealing in the solid state have been published elsewhere [13]. 3.2. Solder joint strength Fig. 4 shows the load versus displacement curves for the solder joints in the as-soldered condition tested at room temperature. Because the thickness of solder in the joint was not strictly constant in all samples, stress–strain curves are not compared. Note that the curves in Fig. 4 are almost linear to failure, indicating the failure of the joints mostly occurred within the elastic strains or with very little plastic strains. Annealed joints also showed similar behaviour. Any plastic deformation must have been restricted to the thin solder between the two Cu blocks. The values of fracture strength (average of maximum stress) obtained for different solders are given in Table 2 for each annealing time, along with their corresponding total IMC layer ( + ε) thickness. A graphical plot of fracture strength against annealing time is shown in Fig. 5. Error bars are significantly large to clearly identify any trends in the graphs. Fracture strengths of Cu/100Sn joints show the widest scatter and no trend could be observed. Strengths of Cu/82Sn18Pb joints show less scatter compared to Cu/100Sn joints. Those of other two solders appear to mildly decrease with annealing time. The large scatter in fracture strength of Cu/100Sn joints could be attributed to the morphology of the IMC layer. The undulating, non-planar -phase/solder interface could have led to the large scatter in the strength of the 100Sn solder joints if the fracture had propagated along
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Fig. 2. Sectional view showing the morphology of the -phase layer at the interface in the as-soldered joints: (a) 100Sn; (b) 82Sn18Pb; (c) 63Sn37Pb and (d) 27Sn73Pb. Labelling is done only in (a) to show the location of the solder, the Cu and the intermetallic. Orientations of the samples in (b–d) are the same as in (a).
this interface. In other solders, the IMC morphology is more planar, as evident from Fig. 3, resulting in less scatter. Several workers have reported that the increase in IMC layer thickness decreases the joint strength [14,15]. For example, Quan et al. [14] reported a 50% decrease in tensile strength of Cu/63Sn37Pb and Cu/5Sn95Pb joints when the total IMC layer thickness increased from an initial 4 m to final thickness of 15 and 11 m, respectively. Our results show a lower decrease in strength of 18.4 and 22% with increase in IMC layer thickness in Cu/63Sn37Pb and Cu/27Sn73Pb joints, respectively. Fig. 6 shows a plot of fracture strength against the Pb-content of the solder at different annealing times. It is evident that the joint strength reaches a maximum at the eutectic composition of 63Sn37Pb. The 100Sn solder has the lowest strength. It is known that the tensile strength of the bulk Sn–Pb solders increases with Pb-content until the eutectic composition (63Sn37Pb), and thereafter decreases with further increase in Pb-content. For example, the tensile strength of 5Sn95Pb is reported to be about 29 MPa, that of 63Sn37Pb is 46 MPa and that of 100Sn is 12 MPa at room temperature [16]. Our results show a similar trend in the
variation of joint strength with Pb-content of the solder. The authors in Ref. [16] did not mention the strain rate at which they performed the tensile testing. Hence we could only compare trends in the effect of Pb-content on the strength, not the actual strengths of the joints. Moreover, there are more potential failure paths in the solder joints than the bulk solders which can only fail through obviously the bulk solder. 3.3. Fracture surface analysis In the uniaxial tension, solder joints must fail at the weakest plane in the three constituents namely, the solder, the IMC or the Cu bar, or at one of the interfaces between these constituents. An examination of the fracture path will thus give an insight to the relative strengths of the constituents and interfaces. As the strength of Cu is considerably larger than those of the others, none of the samples failed through Cu. Fracture path was observed to go through various constituents and locations depending on the annealing time, solder composition and the nature of the interfaces. These will be discussed below.
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Fig. 3. Sectional view showing -phase and ε-phase layers at the interface in joints annealed at 175 ◦ C for 16 days: (a) 100Sn; (b) 82Sn18Pb; (c) 63Sn37Pb and (d) 27Sn73Pb. Labelling is done only in (a) to show the location of the solder, the Cu and the intermetallics. Orientations of the samples in (b–d) are the same as in (a).
3.3.1. The high-Pb solder 27Sn73Pb Fracture was found to occur totally through the bulk solder, in a ductile manner in the high-Pb solder 27Sn73Pb under some processing conditions. Fig. 7 shows the fracture surface and a sectional view of a Cu/27Sn73Pb joint in the as-soldered condition. The feature of microvoid coalescence mechanism typical of ductile failure is apparent in Fig. 7a, while the sectional view of Fig. 7b shows that the fracture path is through the bulk solder. After annealing for 1 day
at 175 ◦ C, the fracture looked similar to that in Fig. 7. The samples annealed for 16 days at 175 ◦ C, however, showed a faceted appearance with no sign of ductility, indicating a different fracture mechanism and path. An example of the fracture surface at this condition is shown in Fig. 8. 3.3.2. The pure Sn solder 100Sn The 100Sn solder is also a weak material. Then, a ductile failure through the solder could be expected in the
Table 2 Fracture strengths and total IMC layer thicknesses for each solder composition and annealing time Time (days)
0 1 3 7 16
Pure Sn
82Sn18Pb
63Sn37Pb
27Sn73Pb
Total IMC layer thickness (m)
Maximum fracture strength (MPa)
Total IMC layer thickness (m)
Maximum fracture strength (MPa)
Total IMC layer thickness (m)
Maximum fracture strength (MPa)
Total IMC layer thickness (m)
Maximum fracture strength (MPa)
5.37 7.63 8.27 9.93 11.81
34.9 44.6 30.0 60.2 34.8
4.22 6.59 8.42 10.08 12.24
65.7 77.1 76.6 66.4 73.4
2.47 4.60 6.13 8.27 11.29
85.8 84.0 85.8 78.8 70.0
2.81 5.08 6.12 7.55 10.38
80.1 65.8 63.9 60.8 62.5
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Fig. 4. Load vs. displacement curves of solder joints made with the four solder in the as-soldered condition. The curve for 82Sn18Pb solder is hidden by the curve for 63Sn37Pb solder, as arrowed in the figure.
Cu/100Sn joints too. Fig. 9 shows the fracture surfaces in Cu/100Sn joints in the as-soldered, 1 and 16 days annealed conditions. The as-soldered sample (Fig. 9a) and 1 day annealed sample (Fig. 9b) show typical ductile failure features in most parts of the fracture but some faceting is evident. The 16 days annealed sample (Fig. 9c) shows a faceted surface similar to that in Fig. 8 for Cu/27Sn73Pb joint annealed for 16 days. The faceted fracture dominant in Fig. 9c is evident at the base of the circular microvoid features in Fig. 9b. A sectional view of a fractured sample of the joint annealed for 1 day (Fig. 10a) clearly shows that, although the fracture path is near the -phase/solder interface, in
Fig. 5. Plot of fracture strength against the annealing time for the joints made of all four solders.
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Fig. 6. Plot of fracture strength against the Pb-content of the solder at various annealing durations.
order to remain perpendicular to the tensile stress, it has propagated through the solder in some areas and through the -phase in others. Thus, the faceted features evident at the base of the microvoids in Fig. 9b are probably the cleaved -phase crystals when the fracture propagates. When the fracture is through the solder, it fails by a ductile, microvoid coalescence mechanism. Even in the as-soldered joint, some -phase crystals were identified at the base of the microvoids. A sectional view of the 16 days annealed joint, shown in Fig. 10b, shows that the fracture path is predominantly through the -phase layer, but is near the -phase/ solder interface. Thus, in 100Sn solder joints, the fracture occurs in the vicinity of the -phase/solder interface in all conditions. The uneven -phase/solder interface typical in these alloys (see Fig. 2) may have nucleated the microvoids during the tensile test, encouraging failure to remain near the interface. However, if the fracture is to occur exactly on the interface, it will have to change direction abruptly and frequently to propagate as the interface is uneven, which would be energetically unfavourable. Fracture through the -phase crystals appears to have been by cleavage, giving the faceted appearance. This is possible since it was previously shown that the -phase grains in the IMC layers have a preferred crystallographic orientation [17]. Then cleavage planes in neighbouring -phase grains could be nearly parallel for the fracture to propagate easily through the IMC. The relative area fractions of the solder and IMC on the fracture surface will vary from sample to sample as the interface morphology would not be identical even if the external processing conditions are nominally the same. This may have caused the high scatter in the fracture strength in joints with 100Sn solder. When annealed, the IMC layer grows and becomes more planar reducing the scatter in the fracture strength.
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Fig. 7. Fracture in a Cu/27Sn73Pb joint in the as-soldered condition: (a) the fracture surface (b) a sectional view of one side. The fracture (arrowed) is through the bulk solder away from the interface.
3.3.3. The low-Pb solder 82Sn18Pb and eutectic solder 63Sn37Pb The fracture surfaces of the as-soldered joints with these solders are shown in Fig. 11. Some ductility is evident. Some -phase grains are visible on the fracture surfaces. The exposure of -phase grains could be because of -phase/solder
Fig. 9. Fracture surfaces in Cu/100Sn joints: (a) as-soldered; (b) annealed for 1 day at 175 ◦ C and (c) annealed for 16 days at 175 ◦ C.
Fig. 8. Fracture surface of a Cu/27Sn73Pb joint in annealed at 175 ◦ C for 16 days.
interface decohesion caused by microvoid formation and growth. Thus, the fracture has occurred in the vicinity of the -phase/solder interface. The fracture surfaces and the corresponding sectional views of samples annealed for 1 and 16 days are shown in Fig. 12. It is clear that the fracture had occurred through the IMC layer exposing the cleaved -phase grains. The sectional view shows that the fracture occurred through the -phase layer, but near the -phase/ε-phase interface. In the as-soldered condition, fracture occurred near the
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Fig. 10. Sectional view of a fracture surface in a Cu/100Sn joint annealed at 175 ◦ C for (a) 1 day and (b) 16 days. Fracture path is arrowed. Intermetallics appear grey. Cu is not visible in (a) as it is in the other half of the fractured sample. In (b) one Cu piece and its interface with solder are visible but the fracture is at the interface with the other Cu piece that is not in this half of the sample.
-phase/solder interface. When annealed for 16 days it appeared to have transferred to the -phase/ε-phase interface.
4. Discussion Isothermal annealing simultaneously coarsens the microstructure of the solder and increases the IMC layer thickness. Coarsening of microstructure is generally expected to decreases the tensile strength. However, according to a study by Takemoto et al. [18] in Sn–Pb alloys, the microstructure coarsening appears to have little effect on its tensile strength. For example, in eutectic 63Sn37Pb solder, they reported a tensile strength of 44.4 MPa in an as-cast sample which, when annealed at 120 ◦ C for 7 and 21 days, decreased slightly to 42.2 and 42.5 MPa, respectively. Thus, the tensile strength of the bulk solder appears to remain constant after a slight initial decrease. However, we cannot attribute the decrease in strength observed on annealing our joints to the solder microstructure coarsening as we have shown that a transition occurs in the fracture path and
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Fig. 11. Fracture surface of the (a) Cu/82Sn18Pb and (b) Cu/63Sn37Pb joints in the as-soldered condition.
mode. As evident from Table 2, the IMC layer grew to 10 m or higher in all joints when annealed at 175 ◦ C for 16 days. The fracture also occurred through the -phase layer in these samples with no features of ductility. Then, strengths obtained in these samples must reflect the strength of the IMC layer irrespective of the solder composition. This appears to be true since the strengths of all 16 days annealed samples appear to approach a similar value in the range 62–73 MPa, except in 100Sn (Fig. 5). The reason for the different behaviour of the Cu/100Sn could be the rough morphology of the -phase/solder interface. In comparison, the as-soldered samples of 100Sn, 82Sn18Pb and 63Sn37Pb solder joints generally fractured near the -phase/solder interface and the joint strength increased with increasing Pb-content (Fig. 6). For example, the fracture strength of Cu/63Sn37Pb (85.8 MPa) is higher than that of Cu/82Sn18Pb joint (65.7 MPa) although their respective fracture surfaces appear similar. A contributory factor for this difference could be the cohesive strength of the interface. It is reported that the surface tension of molten pure Pb (0.47 N/m K) is lower than that of pure Sn (0.55 N/m K) at their respective melting temperatures [19]. This implies that molten Pb will wet solids more easily than molten Sn. If we consider wetting as an indicator of the
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Fig. 12. Fracture in a Cu/82Sn18Pb joint annealed at 175 ◦ C for 1 day: (a) the fracture surface (b) a sectional view. Fracture in a Cu/63Sn37Pb joint annealed at 175 ◦ C for 16 days: (c) the fracture surface (d) a sectional view. Sectional views are labelled to show the fracture path, the solder and the intermetallics.
cohesive strength, then the cohesive strength of the -phase/Pb interface should be higher than that of the -phase/Sn interface. As the fraction of -phase/Pb interface increases with increasing Pb-content of the solder, the average cohesive strength of the -phase/solder interface could also increase. The high-Pb solders will then have a high interface cohesive strength and thus the fracture should occur away from the interface as observed in as-soldered Cu/27Sn73Pb joint. The Pb containing solder joints fractured through the -phase after annealing. Annealing thickens the IMC layer and promotes accumulation of Pb at the -phase/solder interface as Sn reacts with Cu to form the IMC. Cohesive strength of the interface could then increase with annealing. Therefore, a combination of thicker IMC layer and stronger -phase/solder interface would favour fracture through -phase grains instead of decohesion at the interface or failure through the bulk solder. 5. Conclusions Tensile testing of Cu/solder joints at room temperature and at constant cross-head speed were performed with four solder compositions 100Sn, 82Sn18Pb, 63Sn37Pb and
27Sn73Pb for annealing times up to 16 days at 175 ◦ C. The fracture strengths and the fracture paths were determined and are related to the microstructure. The following conclusions can be made from this study: 1. The as-soldered joints have a layer of -phase. The morphology and thickness of the -phase varied depending on the solder composition and annealing time. The morphology was uneven in 100Sn solder but became scalloped as the Pb-content increased. 2. The fracture strength of the joint, in general, increased with increasing Pb-content in the solder alloy up to the eutectic composition (63Sn37Pb) and then decreased with further increase in Pb. 3. The fracture strength of Cu/100Sn joints showed a large scatter with annealing time compared to those of the other solders. This could be attributed to the uneven interface between the -phase and the solder in these joints. 4. In the as-soldered condition, fracture occurred near the -phase/solder interface, except in Cu/27Sn73Pb joints. The fracture mechanism in these joints appears to be partly by microvoid coalescence in the bulk solder, partly by cleavage of -phase grains and partly by decohesion at the interface. In Cu/27Sn73Pb joints, however, fracture
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occurred totally by microvoid coalescence through the bulk solder in the as-soldered condition. 5. Annealing appears to transfer the fracture path to the -phase layer. The 16 days annealed samples predominantly failed by cleavage of the -phase grains. The fracture strengths of these samples also appear to converge to a value in a small band irrespective of the solder composition. 6. Increasing Pb-content in the solder could improve the cohesive strength of the -phase/solder interface and contribute to the increase in fracture strength of the joint.
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