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Effects of pad metallization on the low cycle fatigue characteristics of Sn-based solder joints
International Journal of Fatigue 48 (2013) 1–8
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International Journal of Fatigue journal homepage: www.elsevier.com/locate/ijfatigue
Effects of pad metallization on the low cycle fatigue characteristics of Sn-based solder joints K.O. Lee a,⇑,1, Jin Yu a, T.S. Park b,2, S.B. Lee b a b
Department of Materials Science and Engineering, Korea Advanced Institute of Science and Technology, 291 Daehak-ro, Yuseong-gu, Daejeon 305-701, South Korea Department of Mechanical Engineering, Korea Advanced Institute of Science and Technology, 291 Daehak-ro, Yuseong-gu, Daejeon 305-701, South Korea
a r t i c l e
i n f o
Article history: Received 23 March 2012 Received in revised form 6 October 2012 Accepted 1 December 2012 Available online 12 December 2012 Keywords: Pb-free solder Low cycle fatigue Fatigue life prediction Pad metallization
a b s t r a c t The type and growth rate of the interfacial intermetallic compound is noticeably affected by joint metallization, which in turn affects the mechanical reliabilities of the solder joints. Replacing Cu by Ni for the Sn0.7Cu, Sn3.5Ag and Cu doped Sn3.5Ag solder joints increased the fatigue resistance by 20–50%. The noticeable difference of the fatigue resistance appeared with apparent change in the failure loci. The solder joints in the Au/Ni pad metallization had most of the failures through the solder bulk while the ones in the Cu pad metallization had the failure mode through the Cu–Sn IMC layer near the Cu–Sn IMC/Cu interface. The source of the strong influence of the Cu substrate on the poor solder joint fatigue properties appears to be the introduction of a thick layer of Cu6Sn5 IMCs with relatively lower fracture toughness as compared to the Ni–Sn IMC. Fatigue resistance of the Bi-doped Sn3.5Ag solder joints was substantially worse than Sn–0.7Cu, Sn3.5Ag, Cu doped Sn3.5Ag solder joints regardless of the pad metallization type used. Failure occurred mainly through the IMC layer at the solder/pad interface. The decreased fatigue resistance on substituting Bi is accompanied by an increase in the hardness of the solder and Bi segregation at the solder/IMC interface. Ó 2012 Elsevier Ltd. All rights reserved.
1. Introduction Concern over the toxicity of lead has led to the introduction of Pb-free solders in microelectronic [1]. There have been a wide range of efforts to find out the Pb-free solders with the comparable melting temperature, reliability and manufacturability to Pb–Sn solder. Currently, the most widely used Pb-free solder alloys are high-Sn based alloys with small additions of Ag, Cu, Bi, and Zn. Besides searching for new Pb-free alloy compositions and improving the soldering process, it is also equally important to find an appropriate pad metallization compatible to the Pb-free solders. The manufacture of fine-pitch solder joints ordinarily includes interfacial reactions which therefore influence the composition, the microstructure eventually their mechanical properties and reliabilities which are quite different from solder bulk properties [2– 8]. Even with the substantial technical need, there is a lack of published research on the pad metallization effect on the Pb-free solder joint reliability. The fatigue behavior of high-Sn joints is particularly important, since the power cycling in electronic pack-
⇑ Corresponding author. Tel.: +1 480 554 7740. E-mail address: [email protected] (K.O. Lee). Present address: Intel Corporation, Chandler, AZ 85226. Present address: Samsung Electronics Co., LTD, Gyeonggi-Do 443-742, South Korea. 1 2
0142-1123/$ - see front matter Ó 2012 Elsevier Ltd. All rights reserved. http://dx.doi.org/10.1016/j.ijfatigue.2012.12.002
age has the consequence that fatigue is the dominant deformation mechanism [9–13]. The present work was undertaken to study the influence of the substrate metallization on the low-cycle fatigue characteristics of Sn-rich solder joints. Ball grid array solder balls of Sn0.7Cu, Sn3.5Ag, Sn3.5Ag-X (X = 0.75 and 1.5Cu) and Sn3.5Ag-Y (Y = 2.5 and 7.5Bi) were employed to assemble the plastic ball grid array laminated modules having two different kinds of pad metallization; Au/Ni and Cu. Fatigue resistance was analyzed using the Coffin–Manson relation and Morrow’s plastic-energy dissipation model.
2. Experimental procedure The geometry of the fatigue specimens used in this study is shown in Fig. 1. Plastic ball grid array substrates were overlapped into a single-lap shear configuration and joined by a 3 3 pattern of nine solder joints. The solder resist openings with the diameter of 0.594 mm were patterned on the Cu pads using photo-lithography process. The pad metallization was either as-made Cu pad for Cu pad metallization or electroless plated with l lm of Ni coated with 0.05 lm of Au to form the Au/Ni pad metallization. The nominal chemical composition of the Sn-based solder alloys; Sn-0.7Cu, Sn-3.5Ag, Sn3.5Ag-X (X = 0.75 and 1.5Cu) and Sn3.5Ag-Y (Y = 2.5 and 7.5Bi) under investigation are presented
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K.O. Lee et al. / International Journal of Fatigue 48 (2013) 1–8
Stainless steel grip
Fig. 2. Reflow profile.
540μm
PCB
Fig. 1. Schematic diagrams showing the lap shear fatigue fixtures.
Table 1 Chemical composition of Sn based solder alloys (in weight %). Alloy
Sn
Ag
Bi
Sn-3.5Ag Sn-3.5Ag-0.75Cu Sn-3.5Ag-1.5Cu Sn-3.5Ag-2.5Bi Sn-3.5Ag-7.5Bi Sn-0.7Cu
95.7 95.1 94.6 93.7 88.5 99.4
3.61 3.65 3.49 3.57 3.63
–
Cu 0.75 1.48
2.51 7.62 0.57
Fig. 4. A typical load drop with the fatigue cycles in the Sn-3.5Ag-0.75Cu alloy.
in Table 1. The solder balls with the diameter of 0.76 mm were placed between the lands on the substrates, and reflowed to form the solder joints in a programmable oven under nitrogen gas. The reflow temperature profile used is shown in Fig. 2. The time above 220 °C and the peak temperature was 110 s and 260 °C, respectively. The post reflow solder joint stand-off height was 540 lm. The reflowed joint was attached to the stainless steels grip using an epoxy adhesive, cyanoacrylate. The micromechanical testing system used for the fatigue experiment has a servomotor attached to a ball screw driven rail table, and the displacement
between grips were controlled to the accuracy of 50 nm using a computer controlled feedback system and an AC linear variable differential transformer (LVDT) [14]. To increase machine stiffness and reduce out of plane forces, a rigid alignment grip was used to align grips as shown in Fig. 1. In the present work, triangular displacement waveforms with the frequency of 1/30 Hz were applied at room temperature under varying total displacement amplitudes ranging from 6 to 20 lm. A typical displacement and correspond-
Fig. 3. (a) A displacement stroke time with the displacement amplitude of ±15 lm and (b) corresponding force.
K.O. Lee et al. / International Journal of Fatigue 48 (2013) 1–8
ing load profiles are shown in Fig. 3. Note that there are time lags between the displacement and load cycles caused by the timedependent plasticity of the solder. This can also be observed easily at the load–displacement hysteresis under the total displacement control tests as shown in Fig. 4. The load amplitude, the difference between the maximum and minimum loads, decreased with the load cycle, presumably due to the degradation of solder joints and the propagation of fatigue cracks [15,16]. The number of load cycles which gave 50% load drop was here defined as the fatigue life. The parasitic displacement coming from the loading systems which is composed of printed circuit board (PCB), adhesives, and stainless steel grip, was measured using specimens without solder ball attachment only with adhesive between the PCBs, and subsequently subtracted from the total displacement to give the true solder joint deformation. The parasitic displacement was linearly proportional to the load and the displacement to load ratio,
(a) 10μm
10μm
(b) 10μm
10μm
10μm
10μm
(c)
(d)
3
0.0507 lm/N was reproducible within the displacement range used here. The joints were hold for 10 days at room temperature before fatigue tests to stabilize the solder microstructure by age softening [17]. To study changes in the microstructure during fatigue, the solder joints were mounted and mechanically polished up to 0.05 lm alumina slurry, etched with a solution containing HNO3, HCl, and CH3OH, and were examined using an optical microscope (OM) and scanning electron microscope (SEM).
3. Experimental results Fig. 5 shows OM micrographs of the interface and matrix microstructure of Sn-based solder joints in the Au/Ni and Cu pad metallization. As the author has discussed in a previous report [18], the dominant intermetallic of the Sn-based solders in the Cu pad metallization is Cu6Sn5, while in the Au/Ni pad metallization, it is Ni3Sn4 except for the Cu doped Sn3.5Ag solders. The Cu doped Sn3.5Ag solder has (Cu–Ni)6Sn5 IMC even in the Au/Ni pad metallization when the Cu concentration is more than 5 wt.% in the solder which is consistent to the report by Kang et al. [11]. The white region in the figure is the primary b-Sn dendrites, which are surrounded by a dark eutectic phase that is composed of tiny Ag3Sn precipitates embedded in the b-Sn phases. The volume of the b-Sn phase was largest for the Sn0.7Cu alloy in both pad metallization. Replacing Cu by Au/Ni on both substrates made the dendrite arm spacing smaller and primary b-Sn dendrite volume larger. The impact of Au and Ni on the fine dendrite arm size appears to be the introduction of more nucleation sites for the dendrite b-Sn matrix to form and alter the eutectic points, which in turn makes the b-Sn dendrite denser [19]. Fig. 6 shows that Ni pad metallization has a marked effect on suppressing the IMC growth as compared to Cu pad metallization. The average thickness of the IMC formed was about 1–1.5 lm and 1.5–3.5 lm on the Au/Ni and Cu pad metallization, respectively depending upon the solder alloy composition used. It is evident that the effect of the dopants of the solder alloys on the IMC growth kinetics is relatively small on Ni pad metallization. On the other hand, Cu and Bi turned out to enhance the growth of the Cu–Sn IMC layer on the Cu pad metallization. The larger thickness variation of the Cu–Sn IMC formed on the Cu pad metallization is mainly from the scallop shape of the IMC while Ni–Sn interfacial IMC has planar shape.
10μm
10μm
(e) 10μm
10μm
10μm
10μm
(f)
Au/N
Cu
Fig. 5. Optical image of the solder joint microstructure of (a) Sn-3.5Ag (b) Sn-3.5Ag0.75Cu (c) Sn-3.5Ag-1.5Cu (d) Sn-3.5Ag-2.5Bi (e) Sn-3.5Ag-7.5Bi and (f) Sn-0.7Cu on Au/Ni and Cu pad metallization.
Fig. 6. IMC thickness of Pb-free solder joints with Au/Ni and Cu pad metallization.
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Δ T = ± 10μm
Δ T = ± 10μm
Δ T = ± 12μm
Δ T = ± 12μm
(a)
(b)
Fig. 7. Maximum force variances of the solder joints with (a) Au/Ni and (b) Cu pad metallization.
Variations of the maximum force (Fmax) in the force–displacement hysteresis with the displacement cycles at DT strokes of 10 and 12 lm for the six solder alloys on both Au/Ni and Cu pad metallization are shown in Fig. 7. The magnitude of Fmax in the beginning of the load cycle decreased in the order of Sn3.5Ag7.5Bi, Sn3.5Ag2.5Bi, Sn3.5Ag1.5Cu–Sn3.5Ag0.75Cu–Sn3.5Ag, and SnCu
alloys on both pad metallization. Note that pad metallization had little impact on the initial Fmax. However, a significant load drop was observed, implying fast fatigue-crack growth rate (da/dN) during the period when the Au/Ni pad metallization was replaced by Cu in the most solder alloys except the Bi doped Sn3.5Ag solder
6
0.1
Δγp
3
ΔW (J/m )
10
0.01 Ni Cu
Sn-3.5Ag Sn-3.5Ag-0.75Cu Sn-3.5Ag-1.5Cu Sn-3.5Ag-2.5Bi Sn-3.5Ag-7.5Bi Sn-0.7Cu
1E-3
5
10
Ni Cu
Sn-3.5Ag Sn-3.5Ag-0.75Cu Sn-3.5Ag-1.5Cu Sn-3.5Ag-2.5Bi Sn-3.5Ag-7.5Bi Sn-0.7Cu
4
10
100
1000
Nf Fig. 8. Coffin–Manson’s plot relating the fatigue life with the plastic strain range.
10
10
100
1000
Fig. 9. Morrow energy model relating plastic strain energy density with the fatigue lives.
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K.O. Lee et al. / International Journal of Fatigue 48 (2013) 1–8
joints. Drastic decreases in Fmax were noticeably evident for the Bicontaining alloy joints regardless of pad metallization type used. The relation between the fatigue life and plastic strain can be described by the Coffin–Manson’s relation:
Dcp Nnf ¼ c1
ð1Þ
where Dcp is the plastic strain range, Nf is fatigue life, and n and c1 are material constants. The plastic strain (Dcp) was obtained by subtracting the elastic strain (Dce) from the total strain (DcT), which could be found by extrapolating the initial tangent of the unloading curve of the hysteresis. Using the fatigue life defined
Table 2 Material constants of Sn-based solders on Au/Ni and Cu pad metallization. Solder
Sn-3.5Ag
Sn-3.5Ag-0.75Cu
Sn-3.5Ag-1.5Cu
Sn-3.5Ag-2.5Bi
Sn-3.5Ag-7.5Bi
Sn-0.7Cu
Au/Ni
n c1 n c1
0.53 0.56 0.45 0.20
0.74 2.87 0.47 0.27
0.54 0.74 0.30 0.08
0.87 1.05 0.61 0.34
0.90 0.41 0.64 0.17
Cu
0.43 0.40 0.88 4.08
Table 3 Material constants of Sn-based solders on Au/Ni and Cu pad metallization. Solder
Sn-3.5Ag
Sn-3.5Ag-0.75Cu
Sn-3.5Ag-1.5Cu
Sn-3.5Ag-2.5Bi
Sn-3.5Ag-7.5Bi
Sn-0.7Cu
Au/Ni
m c2 m c2
0.64 2.35E + 07 0.63 9.64E + 06
0.79 6.50E + 07 0.78 2.40E + 07
0.87 1.12E + 08 0.42 3.41E + 06
0.84 1.77E + 07 0.81 1.48E + 07
0.82 7.68E + 06 0.76 5.85E + 06
Cu
0.56 1.33E + 07 1.15 3.06E + 08
(a) 200μm
35μm
200μm
35μm
200μm
200μm
35μm
200μm
35μm
200μm
35μm
200μm
35μm
200μm
35μm
35μm
(b) 200μm
35μm
(c)
(d)
(e) 200μm 35μm
(f) 200μ μm
Pad side surface
35μm
Cross section
Au/Ni pad metallization
35μm
200μm
Pad side surface
Cross section
Cu pad metallization
Fig. 10. Fractography of (a) Sn-3.5Ag (b) Sn-3.5Ag-0.75Cu (c) Sn-3.5Ag-1.5Cu (d) Sn-3.5Ag-2.5Bi (e) Sn-3.5Ag-7.5Bi and (f) Sn-0.7Cu on Au/Ni and Cu pad metallization.
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K.O. Lee et al. / International Journal of Fatigue 48 (2013) 1–8
fatigue resistance is less sensitive to the plastic strain range. The material constants of Sn-3.5Ag–Bi alloys had the highest n and c1 among all other solder alloys indicating that this alloys is quite sensitive to the plastic range and worst fatigue resistance. When the Cu pad metallization is used instead of Ni, the material constant, n, became higher except Sn-3.5Ag-0.75Cu and Sn-3.5Ag7.5Bi. On the other hand material constant, c1 became lower. Another model which is commonly used to predict fatigue life is the Morrow energy model [19,20] where the plastic strain energy density (DW) is used instead of the plastic strain:
as the load cycle that gives 50% load drop, a Coffin–Manson type plot was made in Fig. 8, and the material constants n and c1 of Eq. (1) for various Sn-based solders are summarized in Table 2. Pad metallization had a significant effect on the fatigue resistance of the Sn3.5Ag, Sn0.7Cu eutectics, and Sn3.5Ag–Cu ternary alloys, but had a negligible effect on the Bi bearing Sn3.5Ag alloys. Replacing Ni by Cu on surface finish increased the fatigue resistance by roughly 90–190, 40–220, 110–650, 0–220% for the Sn3.5Ag, Sn3.5Ag0.75Cu, Sn3.5Ag1.5Cu, and Sn0.7Cu, respectively depending upon total strain applied. With addition of Bi in the Sn3.5Ag, no significant fatigue resistance difference was seen between two pad metallization studied. The joints on the Au/Ni pad metallization had lower material constants, n than the ones on the Cu pad metallization except Sn-0.7Cu which implies that the
Nm f DW ¼ c 2
ð2Þ
Here, m and c2 are material constants, and the plastic strain energy density, DW was determined from the area of the first a few hyster-
Ni-Sn IMC Solder
Solder Ni-Sn IMC
Ni-P
Ni3P
Ni3P 5μm
Ni-P
5μm
Cu
(a)
(b)
Cu-Sn IMC Solder
Cu
5μm
Cu
Solder
Cu-Sn IMC
(c)
5μm
(d) Solder Cu-Sn IMC
Bi
Cu 5μm
(e) Fig. 11. Back scattered SEM image of cross-section of solder joints (a) Sn-3.5Ag (b) Sn-3.5Ag-7.5Bi on Au/Ni pad metallization (c) Sn-0.7Cu (d) Sn-3.5Ag and (e) Sn-3.5Ag-7.5Bi on Cu pad metallization.
Solder
Solder
Ni3Sn4 Ni3P Ni(P)
(a)
10μm
Cu6Sn5 Cu
10μm
(b)
Fig. 12. The IMC thickness and morphology of solder joint with (a) Au/Ni and (b) Cu pad metallization.
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K.O. Lee et al. / International Journal of Fatigue 48 (2013) 1–8
100
Peak load (N)
80
60
40
Ni
Cu
20
Sn-3.5Ag Sn-3.5Ag-1.5Cu Sn-3.5Ag-7.5Bi Sn-0.7Cu
0 6
8
10
12
14
16
18
20
22
16
18
20
22
ΔT
(a) 0.05
Ni 0.04
Cu
Sn3.5Ag Sn3.5Ag1.5Cu Sn3.5Ag7.5Bi Sn0.7Cu
0.03 p
esis loops. Prediction of the Morrow model is presented in Fig. 9, which is more or less similar to those of the Coffin–Manson equation. A major difference is that the Morrow model predicted that the fatigue resistance of the Bi doped Sn3.5Ag alloys was very close between two pad metallization types used. Material constants of Eq. (2) are listed in Table 3. It can be seen that the Sn0.7Cu on the Cu pad metallization had the highest m and c2 among all the solder joints studied, which implies the best fatigue resistance. The fracture surfaces of lead-free solder joints after fatigue failure are shown in Fig. 10. Pad metallization had marked effect on the failure locus. For the Sn-3.5Ag, Cu doped Sn-3.5Ag and Sn0.7Cu alloys with the Au/Ni pad metallization, cracks were initiated at the solder wedge and then propagated into the solder matrix by inter-granular while for the solder joints on the Cu pad metallization, cracks initiated at the surface of the Cu-IMC and solder interface, then followed the Cu-IMC/solder interface. However, fatigue failure locus of the Bi-containing Sn3.5Ag alloys were always through the solder/pad metallization interface regardless of the pad metallization type used. Detailed observations of the solder/pad metallization interface areas were performed using a backscattered SEM to reveal a salient cracking feature as shown in Fig. 11. For the Sn3.5Ag on the Ni pad metallization which may be considered as a group having a fatigue failure through solder matrix, micro cracks were also observed interior of the interfacial IMC, but did not evolve as a main crack leading to a total fracture probably due to limited crack density. On the other hand, the Sn3.5Ag on the Cu pad metallization had a high density of cracks in the interfacial Cu–Sn IMC. The cracks eventually coalesced and resulted in macroscopic failure. The higher density of the interfacial IMC cracks for the joints on the Cu pad metallization may be attributed to the lower fracture toughness of the Cu6Sn5 IMC as compared to the Ni3Sn4 IMC, formed on the Au/ Ni pad metallization. Gosh reported that the facture toughness of Cu6Sn5 and Ni3Sn4 is 2.8 MPa m1/2 and 4.2 MPa m1/2, respectively [21]. In addition to the lower fracture toughness of the IMC formed on the Cu pad metallization, other factors served to increase the crack density are thicker Cu–Sn IMC layer as compared to the Ni–Sn IMC layer. The fatigue life of a joint can be reduced by having excess IMCs at the joint interfaces [22]. Fig. 12 shows that the IMC of the solder joints with the Cu pad metallization was about three times thicker as compared the ones formed on the Au/Ni pad metallization. For the Bi-containing alloy joints on the Au/Ni pad metallization, the fatigue crack propagated along the Ni3Sn4/Ni3P interface, where the thin Ni3P layers were a byproduct of the reaction-assisted self-crystallization of amorphous Ni(P) during reflow [23]. It is known that the self-crystallization of Ni(P) and concomitant formation of Ni3Sn4 produce high residual stress in the layer which is often a cause of premature brittle failure [24]. The reason why the fatigue crack followed the different interface only in the Bicontaining alloys on the Au/Ni pad metallization remains unclear at the moment. It appears that the high stiffness of solder matrix reinforced with Bi solid solution hardening tends to increase the propensity for brittle cracking [25]. Fig. 13a shows the peak load at the first cycle with respect to the total displacement. The Bi containing alloys had much higher peak load than other alloys, which implies that the stress on the IMC for the Sn3.5AgBi alloys was much higher than the other alloys. Similarly, to assess how much strain the solder matrix took rather than the IMC, the plastic strain (cp) was plotted as a function of total displacement in Fig. 13b, and the Bi containing alloys had the smaller cp due to high stiffness of the material which indicates that the IMC had to take most of the applied strain during fatigue test. On the other hand, the Sn-0.7Cu alloy had the lowest peak load and the largest cp resulting in good fatigue resistance on the both pad metallization types used. Therefore, it appears that soft matrix and superior plastic flow capacity
0.02
0.01
0.00 6
8
10
12
14
ΔT
(b) Fig. 13. (a) Peak loads of Pb-free solders (b) plastic strain range (cP) as function of the total displacement range (DT).
are important criteria to consider in the design of fatigue resistant materials on the Au/Ni pad metallization. A parallel work on the fatigue behavior of Sn-3.5Ag–Bi alloys on Cu showed crack propagation mainly on solder/Cu6Sn5 interfaces, suggesting that the failure location was affected by the type of pad metallization. It appears the Bi segregation which often observed on the solder/Cu6Sn5 interface for the Bi-containing alloys leads to delamination of solder/Cu6Sn5 interface [26]. 4. Conclusion Pad metallization had a marked effect on the fatigue resistance of the Sn3.5Ag, Sn0.7Cu eutectics, and Sn3.5Ag–Cu ternary alloys, but had a negligible impact on the Bi bearing Sn3.5Ag alloys. When Au/Ni pad metallization was replaced by Cu, the fatigue resistance of the Sn3.5Ag, Sn0.7Cu eutectics, and Sn3.5Ag–Cu ternary alloys was degraded by about 20–50%. The apparent reasons for the decrease in fatigue resistance when Ni metallization was replaced by Cu are thicker Cu–Sn interfacial IMC with lower fracture toughness as compared to the Ni–Sn interfacial IMC. Fatigue resistance of the Bi-doped Sn3.5Ag solder joints was substantially worse than Sn-0.7Cu, Sn3.5Ag, Cu doped Sn3.5Ag solder joints regardless of
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K.O. Lee et al. / International Journal of Fatigue 48 (2013) 1–8
the pad metallization type used. The decreased fatigue resistance on substituting Bi seems to be accompanied by an increase in the hardness of the solder and Bi segregation at the solder/IMC interface. It appears that soft solder matrix and superior plastic flow capacity with robust interfacial IMC are the pre-requisites for high low cycle fatigue resistance. Acknowledgements This work was supported by Center for Electronic Packaging Materials which is an engineering research center designated by the Korea Science and Engineering Foundation. S.B. Lee acknowledges the support from National Research Foundation of Korea under Grant No. 2011-0027669. References [1] European parliament, Proposal for a directive of the European parliament and of the council on waste electrical and electronic equipment and on the restriction of the use of certain hazardous substances in electrical and electronic equipment, COM, 347 (2000). [2] Kariya Y, Otsuka M. J Electron Mater 1998;27(11):1229. [3] Kanchanomai C, Miyahita Y, Mutoh Y. Int J Fatigue 2002;24:671.
[4] Liang J, Gollhardt N, Lee PS, Schrodeder SA, Morris WL. Fatigue Fract Eng Mater Struct 1996;19(11):1401. [5] Kanchanomai C, Yamamoto S, Miyahsita Y, Mutoh Y, Mcevily AJ. Int J Fatigue 2002;24:57. [6] Wen S, Keer LM, Mavoori H. J Electron Mater 2001;30(9):1190–6. [7] Zhao J, Mutoh Y, Miyashita Y, Mannan SL. J Electron Mater 2002;31(8):879. [8] Song HG, Morris Jr JW, Hua F. JOM 2002;6:30. [9] Verma K, Han B. J Electron Packaging 2000;122:227. [10] Lee WW, Nguyen LT, Selvaduray GS. Microelectron Reliab 2000;40:231. [11] Kang SK, Sarkhel AK. J Electron Mater 1994;23(8):701. [12] Lee SM, Stone DS. Scripta Metall et Mater 1994;30(9):1213. [13] Kanchanomai C, Miyashita Y, Mutoch Y. Int J Fatigue 2002;24:987. [14] Park TS, Lee SB. 52th ECTC 2002;979. [15] Kariya Y, Otsuka M. J Electron Mater 1998;27(11):1229. [16] Lee KO, Yu Jin, Park TS, Lee SB. J Electron Mater 2004;33(4):249. [17] Lin JK, Silva AD, Frear D, Guo Y, Jang JW, Li L, et al. 51st ECTC 2001;455. [18] Lee K-O, Morris Jr JW, Hua F. Metall Mater Trans A 2010;41:1805. [19] Kang SK, Choi WK, Shin DY, Lauro P, Henderson DW, Gosselin T, et al. 52nd ECTC 2002;146. [20] Solomon HD et al. J Electro Pack 1996;118:67. [21] Ghosh G. J Mater Res 2004;19(5):1439. [22] Chiou BS, Change JH, Du JG. IEEE Trans Comput Packag Manufact Technol 1995;18:537. [23] Jang JW, Frear DR, Lee TY, Tu KN. J Appl Phys 2000;88:6359. [24] Song JY, Yu J. Thin Solid Films 2002;415:167. [25] Choi S, Subramanian KN, Lucas JP, Bieler TR. J Electron Mater 2000;29(10):1249. [26] Huh JY, Han S, Yong Park CY. Metal Mater Int 2004;10(2 12).
Contents lists available at SciVerse ScienceDirect
International Journal of Fatigue journal homepage: www.elsevier.com/locate/ijfatigue
Effects of pad metallization on the low cycle fatigue characteristics of Sn-based solder joints K.O. Lee a,⇑,1, Jin Yu a, T.S. Park b,2, S.B. Lee b a b
Department of Materials Science and Engineering, Korea Advanced Institute of Science and Technology, 291 Daehak-ro, Yuseong-gu, Daejeon 305-701, South Korea Department of Mechanical Engineering, Korea Advanced Institute of Science and Technology, 291 Daehak-ro, Yuseong-gu, Daejeon 305-701, South Korea
a r t i c l e
i n f o
Article history: Received 23 March 2012 Received in revised form 6 October 2012 Accepted 1 December 2012 Available online 12 December 2012 Keywords: Pb-free solder Low cycle fatigue Fatigue life prediction Pad metallization
a b s t r a c t The type and growth rate of the interfacial intermetallic compound is noticeably affected by joint metallization, which in turn affects the mechanical reliabilities of the solder joints. Replacing Cu by Ni for the Sn0.7Cu, Sn3.5Ag and Cu doped Sn3.5Ag solder joints increased the fatigue resistance by 20–50%. The noticeable difference of the fatigue resistance appeared with apparent change in the failure loci. The solder joints in the Au/Ni pad metallization had most of the failures through the solder bulk while the ones in the Cu pad metallization had the failure mode through the Cu–Sn IMC layer near the Cu–Sn IMC/Cu interface. The source of the strong influence of the Cu substrate on the poor solder joint fatigue properties appears to be the introduction of a thick layer of Cu6Sn5 IMCs with relatively lower fracture toughness as compared to the Ni–Sn IMC. Fatigue resistance of the Bi-doped Sn3.5Ag solder joints was substantially worse than Sn–0.7Cu, Sn3.5Ag, Cu doped Sn3.5Ag solder joints regardless of the pad metallization type used. Failure occurred mainly through the IMC layer at the solder/pad interface. The decreased fatigue resistance on substituting Bi is accompanied by an increase in the hardness of the solder and Bi segregation at the solder/IMC interface. Ó 2012 Elsevier Ltd. All rights reserved.
1. Introduction Concern over the toxicity of lead has led to the introduction of Pb-free solders in microelectronic [1]. There have been a wide range of efforts to find out the Pb-free solders with the comparable melting temperature, reliability and manufacturability to Pb–Sn solder. Currently, the most widely used Pb-free solder alloys are high-Sn based alloys with small additions of Ag, Cu, Bi, and Zn. Besides searching for new Pb-free alloy compositions and improving the soldering process, it is also equally important to find an appropriate pad metallization compatible to the Pb-free solders. The manufacture of fine-pitch solder joints ordinarily includes interfacial reactions which therefore influence the composition, the microstructure eventually their mechanical properties and reliabilities which are quite different from solder bulk properties [2– 8]. Even with the substantial technical need, there is a lack of published research on the pad metallization effect on the Pb-free solder joint reliability. The fatigue behavior of high-Sn joints is particularly important, since the power cycling in electronic pack-
⇑ Corresponding author. Tel.: +1 480 554 7740. E-mail address: [email protected] (K.O. Lee). Present address: Intel Corporation, Chandler, AZ 85226. Present address: Samsung Electronics Co., LTD, Gyeonggi-Do 443-742, South Korea. 1 2
0142-1123/$ - see front matter Ó 2012 Elsevier Ltd. All rights reserved. http://dx.doi.org/10.1016/j.ijfatigue.2012.12.002
age has the consequence that fatigue is the dominant deformation mechanism [9–13]. The present work was undertaken to study the influence of the substrate metallization on the low-cycle fatigue characteristics of Sn-rich solder joints. Ball grid array solder balls of Sn0.7Cu, Sn3.5Ag, Sn3.5Ag-X (X = 0.75 and 1.5Cu) and Sn3.5Ag-Y (Y = 2.5 and 7.5Bi) were employed to assemble the plastic ball grid array laminated modules having two different kinds of pad metallization; Au/Ni and Cu. Fatigue resistance was analyzed using the Coffin–Manson relation and Morrow’s plastic-energy dissipation model.
2. Experimental procedure The geometry of the fatigue specimens used in this study is shown in Fig. 1. Plastic ball grid array substrates were overlapped into a single-lap shear configuration and joined by a 3 3 pattern of nine solder joints. The solder resist openings with the diameter of 0.594 mm were patterned on the Cu pads using photo-lithography process. The pad metallization was either as-made Cu pad for Cu pad metallization or electroless plated with l lm of Ni coated with 0.05 lm of Au to form the Au/Ni pad metallization. The nominal chemical composition of the Sn-based solder alloys; Sn-0.7Cu, Sn-3.5Ag, Sn3.5Ag-X (X = 0.75 and 1.5Cu) and Sn3.5Ag-Y (Y = 2.5 and 7.5Bi) under investigation are presented
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Stainless steel grip
Fig. 2. Reflow profile.
540μm
PCB
Fig. 1. Schematic diagrams showing the lap shear fatigue fixtures.
Table 1 Chemical composition of Sn based solder alloys (in weight %). Alloy
Sn
Ag
Bi
Sn-3.5Ag Sn-3.5Ag-0.75Cu Sn-3.5Ag-1.5Cu Sn-3.5Ag-2.5Bi Sn-3.5Ag-7.5Bi Sn-0.7Cu
95.7 95.1 94.6 93.7 88.5 99.4
3.61 3.65 3.49 3.57 3.63
–
Cu 0.75 1.48
2.51 7.62 0.57
Fig. 4. A typical load drop with the fatigue cycles in the Sn-3.5Ag-0.75Cu alloy.
in Table 1. The solder balls with the diameter of 0.76 mm were placed between the lands on the substrates, and reflowed to form the solder joints in a programmable oven under nitrogen gas. The reflow temperature profile used is shown in Fig. 2. The time above 220 °C and the peak temperature was 110 s and 260 °C, respectively. The post reflow solder joint stand-off height was 540 lm. The reflowed joint was attached to the stainless steels grip using an epoxy adhesive, cyanoacrylate. The micromechanical testing system used for the fatigue experiment has a servomotor attached to a ball screw driven rail table, and the displacement
between grips were controlled to the accuracy of 50 nm using a computer controlled feedback system and an AC linear variable differential transformer (LVDT) [14]. To increase machine stiffness and reduce out of plane forces, a rigid alignment grip was used to align grips as shown in Fig. 1. In the present work, triangular displacement waveforms with the frequency of 1/30 Hz were applied at room temperature under varying total displacement amplitudes ranging from 6 to 20 lm. A typical displacement and correspond-
Fig. 3. (a) A displacement stroke time with the displacement amplitude of ±15 lm and (b) corresponding force.
K.O. Lee et al. / International Journal of Fatigue 48 (2013) 1–8
ing load profiles are shown in Fig. 3. Note that there are time lags between the displacement and load cycles caused by the timedependent plasticity of the solder. This can also be observed easily at the load–displacement hysteresis under the total displacement control tests as shown in Fig. 4. The load amplitude, the difference between the maximum and minimum loads, decreased with the load cycle, presumably due to the degradation of solder joints and the propagation of fatigue cracks [15,16]. The number of load cycles which gave 50% load drop was here defined as the fatigue life. The parasitic displacement coming from the loading systems which is composed of printed circuit board (PCB), adhesives, and stainless steel grip, was measured using specimens without solder ball attachment only with adhesive between the PCBs, and subsequently subtracted from the total displacement to give the true solder joint deformation. The parasitic displacement was linearly proportional to the load and the displacement to load ratio,
(a) 10μm
10μm
(b) 10μm
10μm
10μm
10μm
(c)
(d)
3
0.0507 lm/N was reproducible within the displacement range used here. The joints were hold for 10 days at room temperature before fatigue tests to stabilize the solder microstructure by age softening [17]. To study changes in the microstructure during fatigue, the solder joints were mounted and mechanically polished up to 0.05 lm alumina slurry, etched with a solution containing HNO3, HCl, and CH3OH, and were examined using an optical microscope (OM) and scanning electron microscope (SEM).
3. Experimental results Fig. 5 shows OM micrographs of the interface and matrix microstructure of Sn-based solder joints in the Au/Ni and Cu pad metallization. As the author has discussed in a previous report [18], the dominant intermetallic of the Sn-based solders in the Cu pad metallization is Cu6Sn5, while in the Au/Ni pad metallization, it is Ni3Sn4 except for the Cu doped Sn3.5Ag solders. The Cu doped Sn3.5Ag solder has (Cu–Ni)6Sn5 IMC even in the Au/Ni pad metallization when the Cu concentration is more than 5 wt.% in the solder which is consistent to the report by Kang et al. [11]. The white region in the figure is the primary b-Sn dendrites, which are surrounded by a dark eutectic phase that is composed of tiny Ag3Sn precipitates embedded in the b-Sn phases. The volume of the b-Sn phase was largest for the Sn0.7Cu alloy in both pad metallization. Replacing Cu by Au/Ni on both substrates made the dendrite arm spacing smaller and primary b-Sn dendrite volume larger. The impact of Au and Ni on the fine dendrite arm size appears to be the introduction of more nucleation sites for the dendrite b-Sn matrix to form and alter the eutectic points, which in turn makes the b-Sn dendrite denser [19]. Fig. 6 shows that Ni pad metallization has a marked effect on suppressing the IMC growth as compared to Cu pad metallization. The average thickness of the IMC formed was about 1–1.5 lm and 1.5–3.5 lm on the Au/Ni and Cu pad metallization, respectively depending upon the solder alloy composition used. It is evident that the effect of the dopants of the solder alloys on the IMC growth kinetics is relatively small on Ni pad metallization. On the other hand, Cu and Bi turned out to enhance the growth of the Cu–Sn IMC layer on the Cu pad metallization. The larger thickness variation of the Cu–Sn IMC formed on the Cu pad metallization is mainly from the scallop shape of the IMC while Ni–Sn interfacial IMC has planar shape.
10μm
10μm
(e) 10μm
10μm
10μm
10μm
(f)
Au/N
Cu
Fig. 5. Optical image of the solder joint microstructure of (a) Sn-3.5Ag (b) Sn-3.5Ag0.75Cu (c) Sn-3.5Ag-1.5Cu (d) Sn-3.5Ag-2.5Bi (e) Sn-3.5Ag-7.5Bi and (f) Sn-0.7Cu on Au/Ni and Cu pad metallization.
Fig. 6. IMC thickness of Pb-free solder joints with Au/Ni and Cu pad metallization.
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Δ T = ± 10μm
Δ T = ± 10μm
Δ T = ± 12μm
Δ T = ± 12μm
(a)
(b)
Fig. 7. Maximum force variances of the solder joints with (a) Au/Ni and (b) Cu pad metallization.
Variations of the maximum force (Fmax) in the force–displacement hysteresis with the displacement cycles at DT strokes of 10 and 12 lm for the six solder alloys on both Au/Ni and Cu pad metallization are shown in Fig. 7. The magnitude of Fmax in the beginning of the load cycle decreased in the order of Sn3.5Ag7.5Bi, Sn3.5Ag2.5Bi, Sn3.5Ag1.5Cu–Sn3.5Ag0.75Cu–Sn3.5Ag, and SnCu
alloys on both pad metallization. Note that pad metallization had little impact on the initial Fmax. However, a significant load drop was observed, implying fast fatigue-crack growth rate (da/dN) during the period when the Au/Ni pad metallization was replaced by Cu in the most solder alloys except the Bi doped Sn3.5Ag solder
6
0.1
Δγp
3
ΔW (J/m )
10
0.01 Ni Cu
Sn-3.5Ag Sn-3.5Ag-0.75Cu Sn-3.5Ag-1.5Cu Sn-3.5Ag-2.5Bi Sn-3.5Ag-7.5Bi Sn-0.7Cu
1E-3
5
10
Ni Cu
Sn-3.5Ag Sn-3.5Ag-0.75Cu Sn-3.5Ag-1.5Cu Sn-3.5Ag-2.5Bi Sn-3.5Ag-7.5Bi Sn-0.7Cu
4
10
100
1000
Nf Fig. 8. Coffin–Manson’s plot relating the fatigue life with the plastic strain range.
10
10
100
1000
Fig. 9. Morrow energy model relating plastic strain energy density with the fatigue lives.
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K.O. Lee et al. / International Journal of Fatigue 48 (2013) 1–8
joints. Drastic decreases in Fmax were noticeably evident for the Bicontaining alloy joints regardless of pad metallization type used. The relation between the fatigue life and plastic strain can be described by the Coffin–Manson’s relation:
Dcp Nnf ¼ c1
ð1Þ
where Dcp is the plastic strain range, Nf is fatigue life, and n and c1 are material constants. The plastic strain (Dcp) was obtained by subtracting the elastic strain (Dce) from the total strain (DcT), which could be found by extrapolating the initial tangent of the unloading curve of the hysteresis. Using the fatigue life defined
Table 2 Material constants of Sn-based solders on Au/Ni and Cu pad metallization. Solder
Sn-3.5Ag
Sn-3.5Ag-0.75Cu
Sn-3.5Ag-1.5Cu
Sn-3.5Ag-2.5Bi
Sn-3.5Ag-7.5Bi
Sn-0.7Cu
Au/Ni
n c1 n c1
0.53 0.56 0.45 0.20
0.74 2.87 0.47 0.27
0.54 0.74 0.30 0.08
0.87 1.05 0.61 0.34
0.90 0.41 0.64 0.17
Cu
0.43 0.40 0.88 4.08
Table 3 Material constants of Sn-based solders on Au/Ni and Cu pad metallization. Solder
Sn-3.5Ag
Sn-3.5Ag-0.75Cu
Sn-3.5Ag-1.5Cu
Sn-3.5Ag-2.5Bi
Sn-3.5Ag-7.5Bi
Sn-0.7Cu
Au/Ni
m c2 m c2
0.64 2.35E + 07 0.63 9.64E + 06
0.79 6.50E + 07 0.78 2.40E + 07
0.87 1.12E + 08 0.42 3.41E + 06
0.84 1.77E + 07 0.81 1.48E + 07
0.82 7.68E + 06 0.76 5.85E + 06
Cu
0.56 1.33E + 07 1.15 3.06E + 08
(a) 200μm
35μm
200μm
35μm
200μm
200μm
35μm
200μm
35μm
200μm
35μm
200μm
35μm
200μm
35μm
35μm
(b) 200μm
35μm
(c)
(d)
(e) 200μm 35μm
(f) 200μ μm
Pad side surface
35μm
Cross section
Au/Ni pad metallization
35μm
200μm
Pad side surface
Cross section
Cu pad metallization
Fig. 10. Fractography of (a) Sn-3.5Ag (b) Sn-3.5Ag-0.75Cu (c) Sn-3.5Ag-1.5Cu (d) Sn-3.5Ag-2.5Bi (e) Sn-3.5Ag-7.5Bi and (f) Sn-0.7Cu on Au/Ni and Cu pad metallization.
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K.O. Lee et al. / International Journal of Fatigue 48 (2013) 1–8
fatigue resistance is less sensitive to the plastic strain range. The material constants of Sn-3.5Ag–Bi alloys had the highest n and c1 among all other solder alloys indicating that this alloys is quite sensitive to the plastic range and worst fatigue resistance. When the Cu pad metallization is used instead of Ni, the material constant, n, became higher except Sn-3.5Ag-0.75Cu and Sn-3.5Ag7.5Bi. On the other hand material constant, c1 became lower. Another model which is commonly used to predict fatigue life is the Morrow energy model [19,20] where the plastic strain energy density (DW) is used instead of the plastic strain:
as the load cycle that gives 50% load drop, a Coffin–Manson type plot was made in Fig. 8, and the material constants n and c1 of Eq. (1) for various Sn-based solders are summarized in Table 2. Pad metallization had a significant effect on the fatigue resistance of the Sn3.5Ag, Sn0.7Cu eutectics, and Sn3.5Ag–Cu ternary alloys, but had a negligible effect on the Bi bearing Sn3.5Ag alloys. Replacing Ni by Cu on surface finish increased the fatigue resistance by roughly 90–190, 40–220, 110–650, 0–220% for the Sn3.5Ag, Sn3.5Ag0.75Cu, Sn3.5Ag1.5Cu, and Sn0.7Cu, respectively depending upon total strain applied. With addition of Bi in the Sn3.5Ag, no significant fatigue resistance difference was seen between two pad metallization studied. The joints on the Au/Ni pad metallization had lower material constants, n than the ones on the Cu pad metallization except Sn-0.7Cu which implies that the
Nm f DW ¼ c 2
ð2Þ
Here, m and c2 are material constants, and the plastic strain energy density, DW was determined from the area of the first a few hyster-
Ni-Sn IMC Solder
Solder Ni-Sn IMC
Ni-P
Ni3P
Ni3P 5μm
Ni-P
5μm
Cu
(a)
(b)
Cu-Sn IMC Solder
Cu
5μm
Cu
Solder
Cu-Sn IMC
(c)
5μm
(d) Solder Cu-Sn IMC
Bi
Cu 5μm
(e) Fig. 11. Back scattered SEM image of cross-section of solder joints (a) Sn-3.5Ag (b) Sn-3.5Ag-7.5Bi on Au/Ni pad metallization (c) Sn-0.7Cu (d) Sn-3.5Ag and (e) Sn-3.5Ag-7.5Bi on Cu pad metallization.
Solder
Solder
Ni3Sn4 Ni3P Ni(P)
(a)
10μm
Cu6Sn5 Cu
10μm
(b)
Fig. 12. The IMC thickness and morphology of solder joint with (a) Au/Ni and (b) Cu pad metallization.
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K.O. Lee et al. / International Journal of Fatigue 48 (2013) 1–8
100
Peak load (N)
80
60
40
Ni
Cu
20
Sn-3.5Ag Sn-3.5Ag-1.5Cu Sn-3.5Ag-7.5Bi Sn-0.7Cu
0 6
8
10
12
14
16
18
20
22
16
18
20
22
ΔT
(a) 0.05
Ni 0.04
Cu
Sn3.5Ag Sn3.5Ag1.5Cu Sn3.5Ag7.5Bi Sn0.7Cu
0.03 p
esis loops. Prediction of the Morrow model is presented in Fig. 9, which is more or less similar to those of the Coffin–Manson equation. A major difference is that the Morrow model predicted that the fatigue resistance of the Bi doped Sn3.5Ag alloys was very close between two pad metallization types used. Material constants of Eq. (2) are listed in Table 3. It can be seen that the Sn0.7Cu on the Cu pad metallization had the highest m and c2 among all the solder joints studied, which implies the best fatigue resistance. The fracture surfaces of lead-free solder joints after fatigue failure are shown in Fig. 10. Pad metallization had marked effect on the failure locus. For the Sn-3.5Ag, Cu doped Sn-3.5Ag and Sn0.7Cu alloys with the Au/Ni pad metallization, cracks were initiated at the solder wedge and then propagated into the solder matrix by inter-granular while for the solder joints on the Cu pad metallization, cracks initiated at the surface of the Cu-IMC and solder interface, then followed the Cu-IMC/solder interface. However, fatigue failure locus of the Bi-containing Sn3.5Ag alloys were always through the solder/pad metallization interface regardless of the pad metallization type used. Detailed observations of the solder/pad metallization interface areas were performed using a backscattered SEM to reveal a salient cracking feature as shown in Fig. 11. For the Sn3.5Ag on the Ni pad metallization which may be considered as a group having a fatigue failure through solder matrix, micro cracks were also observed interior of the interfacial IMC, but did not evolve as a main crack leading to a total fracture probably due to limited crack density. On the other hand, the Sn3.5Ag on the Cu pad metallization had a high density of cracks in the interfacial Cu–Sn IMC. The cracks eventually coalesced and resulted in macroscopic failure. The higher density of the interfacial IMC cracks for the joints on the Cu pad metallization may be attributed to the lower fracture toughness of the Cu6Sn5 IMC as compared to the Ni3Sn4 IMC, formed on the Au/ Ni pad metallization. Gosh reported that the facture toughness of Cu6Sn5 and Ni3Sn4 is 2.8 MPa m1/2 and 4.2 MPa m1/2, respectively [21]. In addition to the lower fracture toughness of the IMC formed on the Cu pad metallization, other factors served to increase the crack density are thicker Cu–Sn IMC layer as compared to the Ni–Sn IMC layer. The fatigue life of a joint can be reduced by having excess IMCs at the joint interfaces [22]. Fig. 12 shows that the IMC of the solder joints with the Cu pad metallization was about three times thicker as compared the ones formed on the Au/Ni pad metallization. For the Bi-containing alloy joints on the Au/Ni pad metallization, the fatigue crack propagated along the Ni3Sn4/Ni3P interface, where the thin Ni3P layers were a byproduct of the reaction-assisted self-crystallization of amorphous Ni(P) during reflow [23]. It is known that the self-crystallization of Ni(P) and concomitant formation of Ni3Sn4 produce high residual stress in the layer which is often a cause of premature brittle failure [24]. The reason why the fatigue crack followed the different interface only in the Bicontaining alloys on the Au/Ni pad metallization remains unclear at the moment. It appears that the high stiffness of solder matrix reinforced with Bi solid solution hardening tends to increase the propensity for brittle cracking [25]. Fig. 13a shows the peak load at the first cycle with respect to the total displacement. The Bi containing alloys had much higher peak load than other alloys, which implies that the stress on the IMC for the Sn3.5AgBi alloys was much higher than the other alloys. Similarly, to assess how much strain the solder matrix took rather than the IMC, the plastic strain (cp) was plotted as a function of total displacement in Fig. 13b, and the Bi containing alloys had the smaller cp due to high stiffness of the material which indicates that the IMC had to take most of the applied strain during fatigue test. On the other hand, the Sn-0.7Cu alloy had the lowest peak load and the largest cp resulting in good fatigue resistance on the both pad metallization types used. Therefore, it appears that soft matrix and superior plastic flow capacity
0.02
0.01
0.00 6
8
10
12
14
ΔT
(b) Fig. 13. (a) Peak loads of Pb-free solders (b) plastic strain range (cP) as function of the total displacement range (DT).
are important criteria to consider in the design of fatigue resistant materials on the Au/Ni pad metallization. A parallel work on the fatigue behavior of Sn-3.5Ag–Bi alloys on Cu showed crack propagation mainly on solder/Cu6Sn5 interfaces, suggesting that the failure location was affected by the type of pad metallization. It appears the Bi segregation which often observed on the solder/Cu6Sn5 interface for the Bi-containing alloys leads to delamination of solder/Cu6Sn5 interface [26]. 4. Conclusion Pad metallization had a marked effect on the fatigue resistance of the Sn3.5Ag, Sn0.7Cu eutectics, and Sn3.5Ag–Cu ternary alloys, but had a negligible impact on the Bi bearing Sn3.5Ag alloys. When Au/Ni pad metallization was replaced by Cu, the fatigue resistance of the Sn3.5Ag, Sn0.7Cu eutectics, and Sn3.5Ag–Cu ternary alloys was degraded by about 20–50%. The apparent reasons for the decrease in fatigue resistance when Ni metallization was replaced by Cu are thicker Cu–Sn interfacial IMC with lower fracture toughness as compared to the Ni–Sn interfacial IMC. Fatigue resistance of the Bi-doped Sn3.5Ag solder joints was substantially worse than Sn-0.7Cu, Sn3.5Ag, Cu doped Sn3.5Ag solder joints regardless of
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the pad metallization type used. The decreased fatigue resistance on substituting Bi seems to be accompanied by an increase in the hardness of the solder and Bi segregation at the solder/IMC interface. It appears that soft solder matrix and superior plastic flow capacity with robust interfacial IMC are the pre-requisites for high low cycle fatigue resistance. Acknowledgements This work was supported by Center for Electronic Packaging Materials which is an engineering research center designated by the Korea Science and Engineering Foundation. S.B. Lee acknowledges the support from National Research Foundation of Korea under Grant No. 2011-0027669. References [1] European parliament, Proposal for a directive of the European parliament and of the council on waste electrical and electronic equipment and on the restriction of the use of certain hazardous substances in electrical and electronic equipment, COM, 347 (2000). [2] Kariya Y, Otsuka M. J Electron Mater 1998;27(11):1229. [3] Kanchanomai C, Miyahita Y, Mutoh Y. Int J Fatigue 2002;24:671.
[4] Liang J, Gollhardt N, Lee PS, Schrodeder SA, Morris WL. Fatigue Fract Eng Mater Struct 1996;19(11):1401. [5] Kanchanomai C, Yamamoto S, Miyahsita Y, Mutoh Y, Mcevily AJ. Int J Fatigue 2002;24:57. [6] Wen S, Keer LM, Mavoori H. J Electron Mater 2001;30(9):1190–6. [7] Zhao J, Mutoh Y, Miyashita Y, Mannan SL. J Electron Mater 2002;31(8):879. [8] Song HG, Morris Jr JW, Hua F. JOM 2002;6:30. [9] Verma K, Han B. J Electron Packaging 2000;122:227. [10] Lee WW, Nguyen LT, Selvaduray GS. Microelectron Reliab 2000;40:231. [11] Kang SK, Sarkhel AK. J Electron Mater 1994;23(8):701. [12] Lee SM, Stone DS. Scripta Metall et Mater 1994;30(9):1213. [13] Kanchanomai C, Miyashita Y, Mutoch Y. Int J Fatigue 2002;24:987. [14] Park TS, Lee SB. 52th ECTC 2002;979. [15] Kariya Y, Otsuka M. J Electron Mater 1998;27(11):1229. [16] Lee KO, Yu Jin, Park TS, Lee SB. J Electron Mater 2004;33(4):249. [17] Lin JK, Silva AD, Frear D, Guo Y, Jang JW, Li L, et al. 51st ECTC 2001;455. [18] Lee K-O, Morris Jr JW, Hua F. Metall Mater Trans A 2010;41:1805. [19] Kang SK, Choi WK, Shin DY, Lauro P, Henderson DW, Gosselin T, et al. 52nd ECTC 2002;146. [20] Solomon HD et al. J Electro Pack 1996;118:67. [21] Ghosh G. J Mater Res 2004;19(5):1439. [22] Chiou BS, Change JH, Du JG. IEEE Trans Comput Packag Manufact Technol 1995;18:537. [23] Jang JW, Frear DR, Lee TY, Tu KN. J Appl Phys 2000;88:6359. [24] Song JY, Yu J. Thin Solid Films 2002;415:167. [25] Choi S, Subramanian KN, Lucas JP, Bieler TR. J Electron Mater 2000;29(10):1249. [26] Huh JY, Han S, Yong Park CY. Metal Mater Int 2004;10(2 12).