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Mechanical characterization of Sn–Ag-based lead-free solders
Microelectronics Reliability 42 (2002) 951–966 www.elsevier.com/locate/microrel
Mechanical characterization of Sn–Ag-based lead-free solders Masazumi Amagai a,*, Masako Watanabe a, Masaki Omiya Kikuo Kishimoto b, Toshikazu Shibuya b a
b,1
,
New Package Development (NPD) Department, Texas Instruments Japan, 4260 Aza Takao, Oaza Kawasaki, Hiji-Machi, Hayami-Gun, Oita-Pref 879-1595, Japan b Department of Mechanical and Science Engineering, Tokyo Institute of Technology, 2-12-1 O-okayama, Meguro-ku, Tokyo 152-8552, Japan Received 7 May 2001; received in revised form 8 January 2002
Abstract Recently, preventing environmental pollutions, lead-free (Pb-free) solders are about to replace tin–lead (Sn–Pb) eutectic solders. However, the mechanical properties of Pb-free solders have not been clarified. Hence, the following study was conducted; first, a rate-dependent plasticity was characterized to represent the inelastic deformation behavior for Sn–Ag-based lead-free solders. The material parameters in a constitutive model were determined in a direct method combining both rate-dependent and rate-independent plastic strains. The constitutive model unifies both rate-dependent creep behavior and rate-independent plastic behavior occurring concurrently at the same time in the solders. Secondly, the strength of solders with a variety of plating materials was studied. Intermetallic compounds (IMC) between solder and electrical pads are formed during reflow process and gradually grow in service. By using the Cu-plates on which Cu or Ni or Ni/Au plating was deposited, the specimens of solder joints were fabricated with Sn–Ag-based lead-free solders. After aging the specimens in an isothermal chamber, tensile tests were performed. From scanning electron microscope (SEM) microscope observation and EDX microprobe analysis, the growth and components of the IMC layer were also examined. Based on the experimental tests, the relations between solder joint strength and the aging period were discussed. Furthermore, the validation of fracture strength of solder joints resulting from the tensile tests was verified with package-mounted board level reliability tests. Ó 2002 Elsevier Science Ltd. All rights reserved.
1. Introduction The tin–lead (Sn–Pb) solder alloy has been widely used as interconnection material in electronic packaging due to its low melting temperatures and good wetting behavior on several substrates such as Cu, Ag, Pd and Au. Recently, due to environmental and health concerns, a variety of new lead-free solders have been developed. Lead-free solders lack the toxicity problems
*
Corresponding author. Tel.: +81-977-73-1729; fax: +81977-73-1582. E-mail addresses: [email protected] (M. Amagai), [email protected] (M. Watanabe), [email protected] (M. Omiya), [email protected] (K. Kishimoto), [email protected] (T. Shibuya). 1 Tel.: +81-3-5734-2501; fax: 81-3-5734-2501.
associated with lead-contained solders. However, unlike lead solders, the recently employed lead-free solders do not have a long history and manufacturing process, and also engineering database have not been established. Especially, mechanical properties of lead-free solder joints have not been clarified. As achieving higher circuit board component densities, package dimensions have been shrinking with decreasing solder bump sizes. Recent trend of significant reduction of solder bump size has resulted in severe requirements for mechanical reliability of the solder joint as compared with previous works [1–8]. Concurrently, the demands for lead free are leading to more rigorous requirements on mechanical characteristics of solder joint reliability. The characterization of a rate-dependent plasticity for lead-free solders has not been seen. General constitutive models are mostly separated as rate-independent plasticity and steady state creep [9,10].
0026-2714/02/$ - see front matter Ó 2002 Elsevier Science Ltd. All rights reserved. PII: S 0 0 2 6 - 2 7 1 4 ( 0 2 ) 0 0 0 1 7 - 3
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However, it is very difficult to separate the plastic strain from creep strain based on the mechanical experimental tests. The parameters in a unified constitutive model can be determined in a direct method combining both ratedependent and rate-independent plastic strains into a viscoplastic strain. Moreover, as decreasing bump sizes, plating materials make the stability of solder joint reliability a grave concern [11,12]. Metallurgical interactions occur between the solders and other materials by means of solid state processes and resulting reaction products (intermetallic compounds, IMC) continue to grow as the solder joint ages. The formation of a thin layer of the IMC is considered desirable for achieving a good metallurgical bond at the interface. However, due to its brittle nature and lattice mismatch, the solder cracks tend to be generated near the IMCs [11,12]. Formation of these IMCs affects the mechanical integrity of solder joints [13–15]. Therefore it is important to study the effect of IMCs on the mechanical properties of the solder joints. In this study, a rate-dependent plasticity and the strength of solders with a variety of plating materials such as Cu, Ni, and Ni/Au platings were characterized for Sn–Ag-based lead-free solders. In the rate-dependent plasticity characterization, the material parameters of a viscoplastic constitutive model for Sn3.5Ag0.75Cu and Sn1.0Ag0.5Cu solder alloy were determined from experimental results. The comparison between the experimental data and model predictions was subsequently conducted. In the work for the strength of solder with a variety of platings, tensile tests were performed for the solder joints after thermal aging. From scanning electron microscope (SEM) observation and EDX microprobe analysis, the growth and components of the IMC layer were also examined. The relations between solder joint strength and aging periods were discussed. Furthermore, the validation of the tensile tests after aging was verified with package-mounted board level reliability tests.
2. Rate-dependent plasticity A rate-dependent plasticity was characterized to represent the inelastic deformation behavior for Sn–Agbased lead-free solders. Commercially Sn3.5Ag0.75Cu and Sn1.0Ag0.5Cu solders were chosen to determine the material parameters of the rate-dependent plasticity. To compare Pb–Sn-based solders, 63Sn37Pb and 62Sn36Pb2Ag were used for experiments. A series of experimental tests for the rate-dependent plasticity was conducted to obtain the available data and subsequently fit the exact values of parameters of the experimental data for a constitutive model. The specimen designs, shown in Fig. 1, were used to evaluate the rate-dependent plasticity behaviors under
Fig. 1. Specimen design of experimental tests.
tensile loading. The bulk solder was cast in air into aluminum molds, and subsequently machined into cylindrical specimens. All tests were conducted in a feedback controlled servo-electric testing machine, and a data acquisition system was used to record load and displacement readings [16]. Tests were carried out under isothermal conditions. The temperature chosen was representative of board level reliability ranging from 25 to 125 °C. All data were obtained in term of stress– strain curves at strain rates, 1, 10 and 100 mm/min. 2.1. Rate-dependent plasticity constitutive model The temperature, stress and state dependencies are incorporated via the Arrhenius term to model steady state strain rate, dep =dt, which can be written as dep Q ð1Þ ¼ A exp ½sinhð1rss Þ1=m dt Rh where A is the pre-exponential factor, Q is the activation energy, R is the Boltzmann constant, h is the absolute temperature, 1 is the multiplier of stress, rss is the steady state stress (the subscript ss denotes quantities relevant only to steady state conditions), and m is the strain rate sensitivity of stress. Eq. (1) has long been used to model steady state behavior in hot working [17]. The simple modification of replacing rss by r=s provides a suitable evolution equation and generates the steady state expression sufficient to capture the major strain hardening characteristics of materials, leading to the following viscoplastic flow equation: 1r i1=m dep Q h ¼ A exp ð2Þ sinh dt Rh s where, r is the equivalent stress, s is the scalar internal variable with dimensions of stress called the deformation resistance. It represents the isotropic resistance to plastic deformation. The evolution equation for s as proposed by Anand [18] can be written as c ¼ cs c
1 sinh1 1
ð3Þ
1 dep exp A dt
Q Rh
m
ð4Þ
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The value of the saturation stress, r , which is determined for each test from the steady state value reached in the stress–strain curves, can be written as c ¼ cs where n 1 dep Q exp s ¼ s A dt Rh
ð5Þ
ð6Þ
where s is the coefficient for deformation resistance saturation value and n is the deformation resistance value. Thus, the value of the steady state stress, r , can be summarized as the following relations: n s 1 dep Q exp r ¼ 1 A dt Rh m
1 dep Q exp ð7Þ sinh1 A dt Rh
Fig. 3. Strain rate as a function of steady state stress for 62Sn36Pb2Ag solder.
Figs. 2–5 show the strain rate as a function of the steady state stress for the 63Sn37Pb, 62Sn36Pb2Ag, Sn3.5Ag0.75Cu and Sn1.0Ag0.5Cu solder, respectively. Based on Eq. (7), the material parameters; A; Q; 1; s; m and n can be fitted to the strain rate versus the steady state stress data. The material parameters for these solders are exhibited in Table 1. 2.2. Plastic strain hardening The strain hardening data, that is the slope of the stress versus plastic strain curve from an isothermal, constant strain rate test, therefore can reflect the evolution of internal variable, s, in Eq. (2). The form of dr=dep which was adopted to represent the hardening behavior of the solders is Fig. 4. Strain rate as a function of steady state stress for Sn3.5Ag0.75Cu solder.
dr r a ¼ ch0 1 p de c
ð8Þ
with c ¼ cs; c
Fig. 2. Strain rate as a function of steady state stress for 63Sn37Pb solder.
c ¼ cs
1 sinh1 1
1 dep exp A dt
ð9Þ
Q Rh
m
ð10Þ
Fig. 6 illustrates an example of experimental data which is the delta stress per the delta plastic strain, dr=dep as a function of the stress per the steady state stress, c=c , for the Sn1.0Ag0.5Cu solder at strain rates, 1, 10 and 100 mm/min. The strain rate sensitivity of hardening, a, was determined using the values of steady state stress, r and ch0 (h0 , hardening constant) which were obtained from a
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Fig. 6. Delta stress per delta plastic strain, dr=dep , as a function of stress per steady state stress, c=c for Sn1.0Ag0.5Cu solder. Fig. 5. Strain rate as a function of steady state stress for Sn1.0Ag0.5Cu solder.
linear least-squares fit of the r and dr=dep data for each test. The parameters of strain hardening factor are summarized for 63Sn37Pb, 62Sn36Pb2Ag, Sn3.5Ag0.75Cu and Sn1.0Ag0.5Cu as exhibited in Table 2. 2.3. Model verification To verify the achieved material parameters of the material constitutive model for solders, the comparison between the experiments and Eq. (2)-based model predictions was conducted using the stress–inelastic strain data. Figs. 7 and 8 illustrate an example of the strain rate behavior of Sn3.5Ag0.75Cu and Sn1.0Ag0.5Cu solder, respectively. A good agreement between the experimental data and the model predictions is achieved for the temperature of the strain rate behavior at 25 °C and the strain rate ranges from 1 to 100 mm/min. This
constitutive model unifies both rate-dependent creep behavior and rate-independent plastic behavior occurring concurrently at the same time in the material. On the other hand, general constitutive models are mostly separated as rate-independent plasticity and steady state creep. However, it is very difficult to separate the plastic strain from creep strain based on the mechanical experimental tests. The achieved parameters in a unified constitutive model can be determined in a direct method combining both rate-dependent and rateindependent plastic strains into a viscoplastic strain and plastic flow term.
3. The strength of solders with a variety of plating materials Due to its brittle nature, the solder cracks tend to be generated near the IMCs. Formation of these IMCs
Table 1 Material parameters for 63Sn37Pb, 62Sn36Pb2Ag, Sn3.5Ag0.75Cu and Sn1.0Ag0.5Cu solder Parameter 1
Q=R (K ) A (s1 ) 1 m s (MPa) n
63Sn37Pb
62Sn36Pb2Ag
Sn3.5Ag0.75Cu
Sn2.0Ag0.5Cu
9400 1:09 107 0.07 0.316 0.998 1:37 105
9400 8:49 106 0.065 0.322 0.990 6:71 104
8400 4:61 106 0.038 0.162 1.04 4:60 103
8400 2:42 107 0.043 0.168 1.005 8:10 104
Table 2 Parameters of strain hardening factor for 63Sn37Pb, 62Sn36Pb2Ag, Sn3.5Ag0.75Cu and Sn1.0Ag0.5Cu solder Parameter
63Sn37Pb
62Sn36n2Ag
Sn3.5Ag0.75Cu
Sn1.0Ag0.5Cu
a h0 (MPa)
1.26 2884
1.33 2951
1.56 3090
1.59 3162
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Fig. 9. Configuration of the test specimen.
Fig. 7. Strain rate behavior of Sn3.5Ag0.75Cu solder.
Fig. 8. Strain rate behavior of Sn1.0Ag0.5Cu solder.
affects the mechanical integrity of solder joints. Therefore it is important to study the effect of IMCs on the mechanical properties of the solder joints. The specimen consists of two copper plates with 0.9 mm thickness, plating layers and solder. The configuration of the test specimen is depicted in Fig. 9. The plating layers were composed of Cu or Ni/Cu or Au/Ni/ Cu structures. Each layer was grown by electroplating.
Cu and Ni layers were 5 lm and Au layer was less than 1 lm thickness, respectively. In this work, Sn–Ag-based solders such as Sn3.5Ag0.75Cu, Sn1.0Ag0.5Cu were used. To compare with conventional solders, Pb–Snbased solders such as 63Sn37Pb, 62Sn36Pb2Ag were also used. Strips of these solders were put between copper plates. Then, infra-red reflow process which is used in standard surface mount technology (SMT) was carried out for the reflow. The reflow temperatures are shown in Table 3. To study the solid states growth kinetics of the IMCs at the solder joints, specimens were placed in the isothermal chamber at 125 °C for 1–113 days. After cooling down for one day at room temperature, tensile tests were performed. The tensile load was applied perpendicular to the solder joint layer with cross-head speed of 0.5 mm/min at 22 °C. After tensile tests, fracture surface was observed by using the SEM. For microstructural observation at the solder joint interface, some of the specimens were sectioned using a low-speed diamond saw and metallographically polished to reveal the interface and the internal microstructure of the solder joints. Microstructural characterization was curried out and the composition of each phase was investigated by EDX microprobe analysis. 3.1. Growth of intermetallic compound layer Cross-sectional morphologies of the intermetallic layers in Cu/62Sn36Pb2Ag solder joints annealed for 0 and 80 days at 125 °C are shown in Fig. 10. Back-scattered
Table 3 Solder types and reflow temperatures Solder types
63Sn37Pb
62Sn26Pb2Ag
Sn3.5Ag0.75Cu
Sn1.0Ag0.5Cu
Melting temperature
183
Reflow temperature Plating metals
240 Cu, Ni, Ni þ Au
170 190 240
216 220 260
219 216 260
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Fig. 10. Microstructures of Cu/62Sn36Pb2Ag solder interface.
electron imaging (BEI) was used to obtain the micrographs, in order to produce better contrast among various layers of material. From left to right, the regions in each of these micrographs are the solidified solder, reaction zone, and Cu layer. IMC region can be seen clearly in the 80 days aged specimen. From EDX microprobe analysis, g-phase Cu6 Sn5 and e-phase Cu3 Sn were found in the reaction zone. Such microstructures of Cu–Sn interfacial IMCs described above are similar to what other researchers have found in bimetallic couples of Sn–Pb solders on copper [19]. The formation of g-phase Cu6 Sn5 intermetallic layers in solder joint during the reflow process arises by interfacial reactions between its constituting species, Sn from the solder and Cu from the copper plate. During isothermal aging, the g-phase Cu6 Sn5 in IMC layer grows by interdiffusion of Cu and Sn and reaction with each other, while the g-phase Cu3 Sn forms and grow by reactions between the Cu substrate and g-phase Cu6 Sn5 layer.
Furthermore, Pb-rich phases were educed adjacent to the intermetallic compounds due to the migration of Sn from solder to the reaction zone. Ni metallization is commonly used as a protective layer on a Cu conductor in electronic devises and circuit fabrications. This layer provides a diffusion barrier to inhibit detrimental growth of Cu–Sn IMCs. Interfacial reaction between 62Sn36Pb2Ag and Ni plating layer is depicted in Fig. 11. Ni/62Sn36Pb2Ag solder joints were aged for 0 and 80 days at 125 °C. From left to right, the regions in each of these micrographs are the solidified solder, reaction zone, Ni layer and Cu layer. IMCs, which were formed between Ni plating and the solder, can be seen clearly in the 80 days aging specimen. The Pb-rich phases were educed adjacent to the IMC layer like Cu plating case. The thickness of IMCs is thinner than that of Cu plating case. From EDX microprobe analysis, IMC layer was composed of Ni–Sn compounds.
Fig. 11. Microstructures of Ni/62Sn36Pb2Ag solder interface.
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Fig. 12. Microstructures of Au–Ni/62Sn36Pb2Ag solder interface.
The Au–Ni–Cu three layer structure is one of the most common solder ball pad finishes for the ball-gridarray (BGA) packages. Interfacial reaction between 62Sn36Pb2Ag and Au–Ni plating layer is shown in Fig. 12. Au þ Ni/62Sn36Pb2Ag solder joints were aged for 0 and 80 days at 125 °C. From left to right, the regions in each of these micrographs are the solidified solder, reaction zone, Ni layer and Cu layer. The Au layer could not be identified here. IMCs region can be seen clearly in the 80 days aged specimen. The Pb-rich phases were educed adjacent to the IMC layer like the previous cases. The thickness of IMCs is almost equal to Ni plating case. From EDX microprobe analysis, IMC layer was composed of Ni–Sn compounds. In Au–Ni plating cases, the Au layer could not be identified from the SEM observation, and also could not be detected by EDX analysis. From the recent study of Kim and Tu [20], Au reacts rapidly with molten Pb–Sn solders to form AuSn4 . This reaction was extremely fast. At 200 °C, the AuSn4 layer could grow to
10 lm thick in 5 s. During the fabrication of the specimen, the Au layer dissolved into the solders very quickly and the Ni layer was exposed. The exposed Ni reacted with the solder to form Ni–Sn compounds at the interface. Therefore, the composition of IMC layer is similar to that of Ni plating case. Figs. 13–15 show the cross-sectional morphologies of Sn3.5Ag0.75Cu solder joints treated by each metallization. The aging periods are for 0 and 80 days at 125 °C. IMCs region can be seen clearly in the 80 days aged specimen. Since this solder is Pb-free type, Pb-rich phases could not be seen at the interface. The thickness of IMC layer is thinner than that of 62Sn36Pb2Ag solder joints for every plating cases. The composition of IMC layer is similar to 62Sn36Pb2Ag solder joints. The thickness of IMC layers for 63Sn37Pb solder and Sn1.0Ag0.5Cu solder with Cu or Ni plating case is depicted in Fig. 16. From this figure, the thickness of IMC layers increases in proportion to the square of aging period. The thickness of IMC layers of Pb-contained
Fig. 13. Microstructures of Cu/Sn3.5Ag0.75Cu solder interface.
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Fig. 14. Microstructures of Ni/Sn3.5Ag0.75Cu solder interface.
Fig. 15. Microstructures of Au þ Ni/Sn3.5Ag0.75Cu solder interface.
solder is larger than that of Pb-free solder. The reaction rate of Ni and Sn is slower than that of Cu and Sn. For the Pb-free solder joints with Ni plating, the IMCs thickness is almost constant during this aging period. IMC layer was formed at reflow process stage, which it is not influenced by the aging periods under 125 °C. This result suggests that, for Pb-free solder, the IMC development can be controlled by using Ni plating pads. 3.2. Tensile strength of the aged solder joints
Fig. 16. Development of the IMC layers.
From the results of the cross-sectional morphologies, several kinds of IMC layers were formed at the interface. The formation of these compounds seems to affect the reliability of the solder joints. In this study, tensile tests were performed for the solder joints after thermal aging and the relation between IMCs growth and the tensile strength were investigated.
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From the observation of the tested specimen, the fracture pattern can be classified under the three types as shown in Fig. 17. Fig. 17(a) shows that the fracture occurs between solder and IMC layer and Fig. 17(b) shows the fracture occurred in the solder itself. Mixed fracture type of these two types is shown in Fig. 17(c). Fig. 18 shows the relation between tensile strength and thermal aging time for Cu plating case. The symbols used in this figure represent the fracture patterns classified in Fig. 17. From this figure, the tensile strengths of the Pb-contained solders are larger than that of the Pb-free solders. As increasing Ag weight (%) in the solder, tensile strength tends to increase. This is the same trend as the results of tensile testing for solders themselves. The tensile strengths decrease with aging period and they become almost constant if the aging period is over 50 days. The decrease rate of the tensile strength for the Pb-free solders is larger than that of Pb-contained solders. For the Pb-contained solders, fracture occurred in the solder at first and after some aging period, fracture pattern changed to the interfacial fracture. On the other hand, fracture pattern of the Pb-free solders was mainly the interfacial fracture.
Fig. 17. Fracture patterns of failed specimen of tensile testing.
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Fig. 18. Effect of aging time on tensile strength for Cu plating case.
The IMC layer development strongly influences the tensile strength of the Pb-free solder joints. Therefore, Cu plating pad finishes are not preferable for the Pb-free solders. Fig. 19 shows the relation between the tensile strength and thermal aging time for Ni plating case. During short aging period, the tensile strengths of the Pb-contained solders are larger than that of the Pb-free solders. However, as the aging time increase, the strengths of the Pb-contained solders decrease and they become almost constant if the aging time period is over 50 days. For the Pb-free solders, the tensile strengths were almost constant during this aging period. After long time aging, the strengths of the Pb-free solder become larger than
Fig. 19. Effect of aging time on tensile strength for Ni plating case.
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Fig. 20. Effect of aging time on tensile strength for Au þ Ni plating case.
that of the Pb-contained solder. For the Pb-contained solders, fracture occurred in the solder at first and after some aging period, fracture pattern tend to be the interfacial fracture. On the other hand, for the Pb-free solders, the fracture occurred in the solders themselves. The interface strengths between solder and Ni are larger than the strengths of the solders themselves. These trends are different from Cu plating case. The effect of aging time on the tensile strength for Au and Ni plating case is depicted in Fig. 20. During short aging period, the tensile strengths of all systems are almost same. As the aging time increase, the strengths of the Pb-contained solders decrease and become constant after 50 days aging. On the other hand, the strengths of the Pb-free solders are almost constant. After short time aging (less than one month), the strengths of the Pb-free solder become larger than those of the Pb-contained
solders. Fracture patterns of the Pb-free solders are mainly the interfacial fracture. The interface strengths between solder and Ni are smaller than the strengths of the solders themselves. These trends are different from Ni plating case. From the SEM and EDX analysis, the IMCs of Au/Ni plating case were similar to Ni plating case. As mentioned before, Au diffused into the solder during reflow process. Ho et al. [21] investigated this phenomenon in detail. At first stage in reflow process, the Au layer reacted very quickly with the solder to form AuSn4 and AuSn2 . AuSn2 is immediately converted to AuSn4 . The growth ratio of AuSn4 and AuSn2 was nearly 1 lm/s. Then, AuSn4 grains began to separate themselves from the Ni layer at the roots of the grains and started to fall into the solder. After AuSn4 left the interface, the Ni layer reacted with Sn to form Ni–Sn compounds. AuSn4 was dispersed into the solder and this distributed AuSn4 contributes to the enhancement of the tensile strength of the solders. Therefore, the interface strengths between solder and Ni become smaller than the strengths of the solders and fracture pattern changes to the interfacial fracture. The reaction rate of Ni and Sn is very slow and the surface condition at the interface is not so different during this aging period. Therefore, thermal aging do not affect the tensile strength of Pb-free solders with Au and Ni pads finishes and this pads finishes is preferable for Pb-free solders. 3.3. Morphology of the fracture surface Fig. 21 shows the fractography of the Cu/62Sn36Pb2Ag solder joints. In this case, fracture occurred at the interface between solder and IMC layer. Before aging, many small dimples can be observed and ductile rupture occurred at the interface. After thermal aging, these dimples become large and the fracture
Fig. 21. Fracture surface of Cu/62Sn36Pb2Ag solder joints.
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surface was filled with these dimples. This is due to the formation of the g-phase Cu6 Sn5 [22]. At first, columnar grains of Cu6 Sn5 are formed on the Cu plates and as the aging times increase, the growth of these columnar grains proceeds through coarsening of the grains. As the dimples become large, the tensile strength decreases and when they become large enough, the tensile strength becomes constant. Fig. 22 shows the fractography of the Ni/62Sn36Pb2Ag solder joints. Before aging, fracture occurred in the solder. The fracture surface is flat and brittle fracture seems to occur due to the plastic constraint. After aging, the fracture type changes to the interfacial fracture. The needle-like Ni3 Sn4 can be observed in the fracture surface. The fracture is supposed to occur inside the IMC layer. Therefore, once the IMC layer formed, the tensile strength depends on the strength of IMC layer itself. The fracture surfaces of the Au þ Ni/62Sn36Pb2Ag solder joints are shown in Fig. 23. In this case, fracture occurred at the interface. Before
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aging, several dimples can be observed and ductile fracture is supposed to occur in this case. After aging, the fracture surface becomes flat and the brittle fracture could occur. The fracture mode is considered to be changed due to thermal aging. Therefore, the tensile strength of this case decreases with aging period. Fig. 24 shows the fractography of Cu/Sn3.5Ag0.75Cu solder joints. In this case, the fracture occurred at the interface. There are many dimples and those dimples become large after thermal aging. These dimples are due to the formation of the Cu6 Sn5 like 62Sn36Pb2Ag case. The fracture surfaces of the Ni/62Sn36Pb2Ag solder joints are shown in Fig. 25. In these cases, the fracture occurred inside the solder before and after aging. The fracture surface is flat and similar to each other. Brittle fracture is supposed to occur in these cases and the tensile strength does not depend on the aging period. Fig. 26 shows the fractography of Au þ Ni/62Sn36Pb2Ag solder joints. In these cases, the fracture occurred at the interface. The
Fig. 22. Fracture surface of Ni/62Sn36Pb2Ag solder joints.
Fig. 23. Fracture surface of Au þ Ni/62Sn36Pb2Ag solder joints.
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Fig. 24. Fracture surface of Cu/Sn3.5Ag0.75Cu solder joints.
Fig. 25. Fracture surface of Ni/Sn3.5Ag0.75Cu solder joints.
Fig. 26. Fracture surface of Au þ Ni/Sn3.5Ag0.75Cu solder joints.
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fracture surface is flat and similar to each other. Brittle fracture is supposed to occur in these cases and the tensile strength does not depend on the aging period.
3.4. Board level reliability tests To verify the variation of specimen samples used in the aging tests at 125 °C, several studies have been conducted using chip scale packages (CSP)-mounted FR-4 board. In this study, the CSP assembles with polyimide-based substrates were tested under thermal cyclic loads. The 10 10 mm2 CSP consists of the silicon chip (6:0 6:0 0:28 mm3 ), an epoxy-based molding compound and 151 solder balls at a 0.5 mm pitch. Two types of solder ball materials were used, 62Sn36Pb2Ag and Sn1.0Ag0.5Cu, respectively. 40 to 125 °C thermal cycles with 5 min ramp and 10 min dwell times were used. The 0.8 mm thick PCB having the Au/Ni or copper pad finish was used. Two types of solder pastes for the PCB pad finish were used, 63Sn37Pb and Sn3.5Ag0.75Cu, respectively. Figs. 27–30 show the three-parameter (3-P) Weibull distribution for the comparison of 62Sn36Pb2Ag versus Sn1.0Ag0.5Cu using the Au–Ni pad finish and 63Sn37Pb solder paste, the Au–Ni pad finish and Sn3.5Ag0.75Cu solder paste, the copper pad finish and 63Sn37Pb solder paste, the copper pad finish and 63Sn37Pb solder paste, respectively. The 3-P Weibull distribution is given by [23]. F ðxÞ ¼ 1 exp
xa Nf ð63:2%Þ a
Fig. 28. Comparison of 62Sn36Pb2Ag versus Sn1.0Ag0.5Cu using the Au–Ni pad finish and Sn3.5Ag0.75Cu solder paste.
Fig. 29. Comparison of 62Sn36Pb2Ag versus Sn1.0Ag0.5Cu using the copper pad finish and 63Sn37Pb solder paste.
b ! ð11Þ
where x is the number of cycles, a is the failure-free life, Nf (63.2%) is the characteristic life and b is the shape parameter which indicates the amount of scatter in the data. The parameters of 3-P Weibull, a; b and 63.2%, are obtained by non-linear least-squares fitting method.
Fig. 30. Comparison of 62Sn36Pb2Ag versus Sn1.0Ag0.5Cu using the copper pad finish and Sn3.5Ag0.75Cu solder paste.
Fig. 27. Comparison of 62Sn36Pb2Ag versus Sn1.0Ag0.5Cu using the Au–Ni pad finish and 63Sn37Pb solder paste.
a is determined by the value of R2 which is a non-linear square fitting parameter closing to 1.0 as much as possible. For the Au–Ni pad finish, Sn1.0Ag0.5Cu solder balls have a factor of 1.16x and 1.52x solder fatigue life (0.1% cumulative fails) of 62Sn36Pb2Ag balls using 63Sn37Pb and Sn3.5Ag0.75Cu solder paste, respectively. On the
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contrary, for the copper pad finish, Sn1.0Ag0.5Cu solder balls have a factor of 0.89xand 0.80x solder fatigue life (0.1% cumulative fails) of 62Sn36Pb2Ag balls using 63Sn37Pb and Sn3.5Ag0.75Cu solder paste, respectively. Based on the results of board level reliability tests, it is found that the Au–Ni pad finish makes Sn1.0Ag0.5Cu a more solder fatigue life than 62Sn36Pb2Ag while the copper pad finish decreases the life of Sn1.0Ag0.5Cu against 62Sn36Pb2Ag. The results of board level reliability tests are similar phenomenon to the results of simple specimen samples used in the aging tests at 125 °C. EPMA was used to determine the density of gold around the interface between the Sn1.0Ag0.5Cu solder and the Ni–Au pad finish. Figs. 31a and b show the density of gold around the interface after the soldering process (a) and few thermal cycles, respectively. After the soldering process, the gold was diffused into the solder. Following few thermal cycles, the gold was reached to the interface once again. It is expected that this phenomenon is due to the following assumption: (1) High stresses are generated
Fig. 32. Strength of 63Sn37Pb as a function of gold weight (%).
near the interface between the solder and the solder pad after cooling down to room temperature and then few thermal cycles, leading to atomic level and solder grain level defects due to dislocation caused by solder creep. (2) The gold, which is diffused into the solder, is moved to the interface once again to repair the atomic level and solder grain level defects. In board level reliability tests, fractures in solder were observed for Sn1.0Ag0.5Cu with the copper pad finish while fractures near the interface were observed for the solder with the Au–Ni pad finish. Fig. 32 reveals that the strength of 63Sn37Pb as a function of gold weight (%) under tensile and shear loads [24,25]. It is obvious that the tensile strength of 63Sn37Pb is increased in range from 0 to 4 wt% while the shear strength is increased in range from 0 to 2 wt%. Without the gold, cracks in the solder are generated for Sn1.0Ag0.5Cu. After the gold diffusion into the solder, due to an increase in the strength of Sn1.0Ag0.5Cu like 63Sn37Pb, and cracks appeared near the interface. For 62Sn36Pb2Ag, fractures near the interface were observed for both the Au/ Ni and copper pad finish. It is found that cracks in the solder are never generated because of high strength (based on the results of the tensile tests in the ratedependent plasticity study). Since cracks near the interface are generated for 62Sn36Pb2Ag, solder fatigue life was determined by microstructural change near the interface induced by the component of solder with a plating material.
4. Conclusion Fig. 31. (a) Density of gold around the interface after the soldering process. (b) Density of gold around the interface after the 125 °C aging process (white color is gold).
A rate-dependent plasticity and the strength of solders with a variety of plating materials such as Cu, Ni,
M. Amagai et al. / Microelectronics Reliability 42 (2002) 951–966
and Ni–Au platings were characterized for Sn–Ag-based lead-free solders. In the rate-dependent plasticity characterization, the material parameters of a viscoplastic constitutive model for Sn3.5Ag0.75Cu and Sn1.0Ag0.5Cu solder alloy were determined from experimental results. This constitutive model unifies both rate-dependent creep behavior and rate-independent plastic behavior occurring concurrently at the same time in the material. A good agreement between the experimental data and the model predictions is achieved for the temperature of the strain rate behavior ranging from 25 to 125 °C and the strain rate ranges from 1 to 100 mm/min. In the study of the strength of solders with a variety of plating materials, tensile tests were performed for the solder joints after thermal aging and the relations between solder joint strength and aging periods were examined. From SEM microscope observation and EDX microprobe analysis, the growth and components of the IMC layer were also examined. For the Pb-contained solders, Cu plating is preferable of the Cu pads finishes, since the strength of tensile strength of Cu plating is larger than the other platings and is not influenced by thermal aging. On the contrary, for the Pb-free solders, Au/Ni plating is preferable of the Cu pads finishes, since the tensile strength of Au/Ni plating is larger than the other plating and is not influenced by thermal aging. In board level reliability tests using 62Sn36Pb2Ag and Sn1.0Ag0.5Cu, the variation of specimen samples used in the aging tests at 125 °C was verified. References [1] Lau John H, Pao Yi-Hsin. Solder joint reliability of BGA, CSP, Flip Chip, and Fine Pitch SMT assemblies. McGrawHill, 1997. p. 1–405. [2] Amagai M. The effect of polymer die attach material on solder joint reliability. In: Proceedings of the ASME Workshop on Mechanical Reliability of Polymeric Materials of IC Devices, Paris, France, November 1998. p. 223–30. [3] Amagai M. Chip scale package solder joint reliability. In: Proceedings of the IEEE 36th International Reliability Physics Symposium, Reno, USA, April 1998. p. 260–8. [4] Darveaux R. Solder joint fatigue life model. In: Design and reliability of solders and solder interconnections. The Minerals, Metals and Materials Society (TMS); 1997. p. 213–8. [5] Amagai M. Chip scale package (CSP) solder joint reliability and modeling. Microelectron Reliab 1999;39:463–77. [6] Amagai M. Chip scale package (CSP) material characterization. Microelectron Reliab 1999;42:573–87. [7] Amagai M. Characterization of molding compound and die attach materials for package warpage and solder joint reliability in chip scale package. In: Proceedings of the ASME InterPack, Hawaii, USA, June 1999. p. 1103–12.
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[8] Amagai M. The effect of solder pad metallization on solder fatigue and crack growth. J Jpn Inst Electr Packaging (JIEP) 2001;4(1):30–6. [9] Mukai M, Kawakami T, Takahashi K, Kishimoto K, Shibuya T. Thermal fatigue life of solder bumps in BGA. JSME Int J A: Solid Mech Mater Eng 1998;41: 260–6. [10] Yu Q, Shiratori M. Thermal fatigue reliability assessment for solder joints of BGA assembly. In: ASME EEPVol 26-1, Advances in electronic packaging, vol. 1, 1999. p. 239–46. [11] Kishimoto K, Masuda K, Omiya M, Shibuya T. Effect of intermetallic compound layer development on mechanical strength of 63Sn–37Pb solder joints. In: Michel B et al., editors. Proceedings of the 3rd International Conference and Poster Exhibition, MicroMaterials, MICROMAT2000, Berlin, Germany, 2000. p. 560– 5. [12] Omiya M, Kishimoto K, Shibuya T, Amagai M. Effect of intermetallic compound layer development on interfacial strength of solder joints. In: ASME EEP-Vol, Advances in electronic packaging, vol. 1, 2001. p. 427–32. [13] Pratt RE, Stromswold EI, Quesnel DJ. Effect of solid-state intermetallic growth on the fracture toughness of Cu/63Sn– 37Pb solder joints. IEEE Trans Components, Packaging, Manufact Technol A 1996;19:134–41. [14] Vianco PT, Stephens JJ, Rejent JA. Intermetallic compound layer development during the solid state thermal aging of 63Sn–37Pb solder/Au–Pt–Pd thick film couples. IEEE Trans Components, Packaging, Manufact Technol A 1997;20:478–90. [15] Chan YC, So ACK, Lai JKL. Growth kinetic studies of Cu–Sn intermetallic compound and its effect on shear strength of LCCC SMT solder joints. Mater Sci Eng B 1998;55:5–13. [16] Wilde J, Becker K, Thoben M, Blum W, Jupitz T, Wang G, Cheng ZN. Rate dependent constitutive relations based on Anand model for 92.5Pb5Sn2.5Ag solder. IEEE Trans Components, Packaging, Manufact Technol B 2000;23: 408–14. [17] Jonas JJ, Sellars CM, Mcg Tegart WJ. Strength and structure under hot working conditions. Metall Rev 1969;14(1). [18] Anand L. Constitutive equations for the rate-dependent deformation of metals at elevated temperatures. Trans ASME, J Eng Mater Technol 1992;104(1):12–7. [19] Choi S, Bieler TR, Lucas JP, Subramanian KN. Characterization of the growth of intermetallic interfacial layers of Sn–Ag and Sn–Pb eutectic solders and their composite solders on Cu substrate during isothermal long-term aging. J Electr Mater 1999;28:1209–15. [20] Kim PG, Tu KN. Fast dissolution and soldering reactions on Au foils. Mater Chem Phys 1998;53:165– 71. [21] Ho CE, Chen YM, Kao CR. Reaction kinetics of solderballs with pads in BGA packages during reflow soldering. J Electr Mater 1999;28:1231–7. [22] Choi WK, Lee HM. Effect of soldering and aging time on interfacial microstructure and growth of intermetallic compounds between Sn–3.5Ag solder alloy and Cu substrate. J Electr Mater 2000;29:1207–13.
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[23] Amagai M. Solder fatigue life of chip scale packages for double-sided surface mount. In: Proceedings of the International Conference on Electronic Packaging, Tokyo, Japan, April 2001. [24] Gangemi L, Pittinato GF. The effect of gold plate on solder joint integrity. MPR 6-117-001, Rocketdyne
Division of North American Aviation, January 1996. p. 466–71. [25] Foster FG. Embrittlement of solder by gold from plated surface. In: Proceedings of the 65th Annual Meeting of ASTM, Special Technical Publication, No. 319, June 1992. p. 13.
Mechanical characterization of Sn–Ag-based lead-free solders Masazumi Amagai a,*, Masako Watanabe a, Masaki Omiya Kikuo Kishimoto b, Toshikazu Shibuya b a
b,1
,
New Package Development (NPD) Department, Texas Instruments Japan, 4260 Aza Takao, Oaza Kawasaki, Hiji-Machi, Hayami-Gun, Oita-Pref 879-1595, Japan b Department of Mechanical and Science Engineering, Tokyo Institute of Technology, 2-12-1 O-okayama, Meguro-ku, Tokyo 152-8552, Japan Received 7 May 2001; received in revised form 8 January 2002
Abstract Recently, preventing environmental pollutions, lead-free (Pb-free) solders are about to replace tin–lead (Sn–Pb) eutectic solders. However, the mechanical properties of Pb-free solders have not been clarified. Hence, the following study was conducted; first, a rate-dependent plasticity was characterized to represent the inelastic deformation behavior for Sn–Ag-based lead-free solders. The material parameters in a constitutive model were determined in a direct method combining both rate-dependent and rate-independent plastic strains. The constitutive model unifies both rate-dependent creep behavior and rate-independent plastic behavior occurring concurrently at the same time in the solders. Secondly, the strength of solders with a variety of plating materials was studied. Intermetallic compounds (IMC) between solder and electrical pads are formed during reflow process and gradually grow in service. By using the Cu-plates on which Cu or Ni or Ni/Au plating was deposited, the specimens of solder joints were fabricated with Sn–Ag-based lead-free solders. After aging the specimens in an isothermal chamber, tensile tests were performed. From scanning electron microscope (SEM) microscope observation and EDX microprobe analysis, the growth and components of the IMC layer were also examined. Based on the experimental tests, the relations between solder joint strength and the aging period were discussed. Furthermore, the validation of fracture strength of solder joints resulting from the tensile tests was verified with package-mounted board level reliability tests. Ó 2002 Elsevier Science Ltd. All rights reserved.
1. Introduction The tin–lead (Sn–Pb) solder alloy has been widely used as interconnection material in electronic packaging due to its low melting temperatures and good wetting behavior on several substrates such as Cu, Ag, Pd and Au. Recently, due to environmental and health concerns, a variety of new lead-free solders have been developed. Lead-free solders lack the toxicity problems
*
Corresponding author. Tel.: +81-977-73-1729; fax: +81977-73-1582. E-mail addresses: [email protected] (M. Amagai), [email protected] (M. Watanabe), [email protected] (M. Omiya), [email protected] (K. Kishimoto), [email protected] (T. Shibuya). 1 Tel.: +81-3-5734-2501; fax: 81-3-5734-2501.
associated with lead-contained solders. However, unlike lead solders, the recently employed lead-free solders do not have a long history and manufacturing process, and also engineering database have not been established. Especially, mechanical properties of lead-free solder joints have not been clarified. As achieving higher circuit board component densities, package dimensions have been shrinking with decreasing solder bump sizes. Recent trend of significant reduction of solder bump size has resulted in severe requirements for mechanical reliability of the solder joint as compared with previous works [1–8]. Concurrently, the demands for lead free are leading to more rigorous requirements on mechanical characteristics of solder joint reliability. The characterization of a rate-dependent plasticity for lead-free solders has not been seen. General constitutive models are mostly separated as rate-independent plasticity and steady state creep [9,10].
0026-2714/02/$ - see front matter Ó 2002 Elsevier Science Ltd. All rights reserved. PII: S 0 0 2 6 - 2 7 1 4 ( 0 2 ) 0 0 0 1 7 - 3
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However, it is very difficult to separate the plastic strain from creep strain based on the mechanical experimental tests. The parameters in a unified constitutive model can be determined in a direct method combining both ratedependent and rate-independent plastic strains into a viscoplastic strain. Moreover, as decreasing bump sizes, plating materials make the stability of solder joint reliability a grave concern [11,12]. Metallurgical interactions occur between the solders and other materials by means of solid state processes and resulting reaction products (intermetallic compounds, IMC) continue to grow as the solder joint ages. The formation of a thin layer of the IMC is considered desirable for achieving a good metallurgical bond at the interface. However, due to its brittle nature and lattice mismatch, the solder cracks tend to be generated near the IMCs [11,12]. Formation of these IMCs affects the mechanical integrity of solder joints [13–15]. Therefore it is important to study the effect of IMCs on the mechanical properties of the solder joints. In this study, a rate-dependent plasticity and the strength of solders with a variety of plating materials such as Cu, Ni, and Ni/Au platings were characterized for Sn–Ag-based lead-free solders. In the rate-dependent plasticity characterization, the material parameters of a viscoplastic constitutive model for Sn3.5Ag0.75Cu and Sn1.0Ag0.5Cu solder alloy were determined from experimental results. The comparison between the experimental data and model predictions was subsequently conducted. In the work for the strength of solder with a variety of platings, tensile tests were performed for the solder joints after thermal aging. From scanning electron microscope (SEM) observation and EDX microprobe analysis, the growth and components of the IMC layer were also examined. The relations between solder joint strength and aging periods were discussed. Furthermore, the validation of the tensile tests after aging was verified with package-mounted board level reliability tests.
2. Rate-dependent plasticity A rate-dependent plasticity was characterized to represent the inelastic deformation behavior for Sn–Agbased lead-free solders. Commercially Sn3.5Ag0.75Cu and Sn1.0Ag0.5Cu solders were chosen to determine the material parameters of the rate-dependent plasticity. To compare Pb–Sn-based solders, 63Sn37Pb and 62Sn36Pb2Ag were used for experiments. A series of experimental tests for the rate-dependent plasticity was conducted to obtain the available data and subsequently fit the exact values of parameters of the experimental data for a constitutive model. The specimen designs, shown in Fig. 1, were used to evaluate the rate-dependent plasticity behaviors under
Fig. 1. Specimen design of experimental tests.
tensile loading. The bulk solder was cast in air into aluminum molds, and subsequently machined into cylindrical specimens. All tests were conducted in a feedback controlled servo-electric testing machine, and a data acquisition system was used to record load and displacement readings [16]. Tests were carried out under isothermal conditions. The temperature chosen was representative of board level reliability ranging from 25 to 125 °C. All data were obtained in term of stress– strain curves at strain rates, 1, 10 and 100 mm/min. 2.1. Rate-dependent plasticity constitutive model The temperature, stress and state dependencies are incorporated via the Arrhenius term to model steady state strain rate, dep =dt, which can be written as dep Q ð1Þ ¼ A exp ½sinhð1rss Þ1=m dt Rh where A is the pre-exponential factor, Q is the activation energy, R is the Boltzmann constant, h is the absolute temperature, 1 is the multiplier of stress, rss is the steady state stress (the subscript ss denotes quantities relevant only to steady state conditions), and m is the strain rate sensitivity of stress. Eq. (1) has long been used to model steady state behavior in hot working [17]. The simple modification of replacing rss by r=s provides a suitable evolution equation and generates the steady state expression sufficient to capture the major strain hardening characteristics of materials, leading to the following viscoplastic flow equation: 1r i1=m dep Q h ¼ A exp ð2Þ sinh dt Rh s where, r is the equivalent stress, s is the scalar internal variable with dimensions of stress called the deformation resistance. It represents the isotropic resistance to plastic deformation. The evolution equation for s as proposed by Anand [18] can be written as c ¼ cs c
1 sinh1 1
ð3Þ
1 dep exp A dt
Q Rh
m
ð4Þ
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The value of the saturation stress, r , which is determined for each test from the steady state value reached in the stress–strain curves, can be written as c ¼ cs where n 1 dep Q exp s ¼ s A dt Rh
ð5Þ
ð6Þ
where s is the coefficient for deformation resistance saturation value and n is the deformation resistance value. Thus, the value of the steady state stress, r , can be summarized as the following relations: n s 1 dep Q exp r ¼ 1 A dt Rh m
1 dep Q exp ð7Þ sinh1 A dt Rh
Fig. 3. Strain rate as a function of steady state stress for 62Sn36Pb2Ag solder.
Figs. 2–5 show the strain rate as a function of the steady state stress for the 63Sn37Pb, 62Sn36Pb2Ag, Sn3.5Ag0.75Cu and Sn1.0Ag0.5Cu solder, respectively. Based on Eq. (7), the material parameters; A; Q; 1; s; m and n can be fitted to the strain rate versus the steady state stress data. The material parameters for these solders are exhibited in Table 1. 2.2. Plastic strain hardening The strain hardening data, that is the slope of the stress versus plastic strain curve from an isothermal, constant strain rate test, therefore can reflect the evolution of internal variable, s, in Eq. (2). The form of dr=dep which was adopted to represent the hardening behavior of the solders is Fig. 4. Strain rate as a function of steady state stress for Sn3.5Ag0.75Cu solder.
dr r a ¼ ch0 1 p de c
ð8Þ
with c ¼ cs; c
Fig. 2. Strain rate as a function of steady state stress for 63Sn37Pb solder.
c ¼ cs
1 sinh1 1
1 dep exp A dt
ð9Þ
Q Rh
m
ð10Þ
Fig. 6 illustrates an example of experimental data which is the delta stress per the delta plastic strain, dr=dep as a function of the stress per the steady state stress, c=c , for the Sn1.0Ag0.5Cu solder at strain rates, 1, 10 and 100 mm/min. The strain rate sensitivity of hardening, a, was determined using the values of steady state stress, r and ch0 (h0 , hardening constant) which were obtained from a
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Fig. 6. Delta stress per delta plastic strain, dr=dep , as a function of stress per steady state stress, c=c for Sn1.0Ag0.5Cu solder. Fig. 5. Strain rate as a function of steady state stress for Sn1.0Ag0.5Cu solder.
linear least-squares fit of the r and dr=dep data for each test. The parameters of strain hardening factor are summarized for 63Sn37Pb, 62Sn36Pb2Ag, Sn3.5Ag0.75Cu and Sn1.0Ag0.5Cu as exhibited in Table 2. 2.3. Model verification To verify the achieved material parameters of the material constitutive model for solders, the comparison between the experiments and Eq. (2)-based model predictions was conducted using the stress–inelastic strain data. Figs. 7 and 8 illustrate an example of the strain rate behavior of Sn3.5Ag0.75Cu and Sn1.0Ag0.5Cu solder, respectively. A good agreement between the experimental data and the model predictions is achieved for the temperature of the strain rate behavior at 25 °C and the strain rate ranges from 1 to 100 mm/min. This
constitutive model unifies both rate-dependent creep behavior and rate-independent plastic behavior occurring concurrently at the same time in the material. On the other hand, general constitutive models are mostly separated as rate-independent plasticity and steady state creep. However, it is very difficult to separate the plastic strain from creep strain based on the mechanical experimental tests. The achieved parameters in a unified constitutive model can be determined in a direct method combining both rate-dependent and rateindependent plastic strains into a viscoplastic strain and plastic flow term.
3. The strength of solders with a variety of plating materials Due to its brittle nature, the solder cracks tend to be generated near the IMCs. Formation of these IMCs
Table 1 Material parameters for 63Sn37Pb, 62Sn36Pb2Ag, Sn3.5Ag0.75Cu and Sn1.0Ag0.5Cu solder Parameter 1
Q=R (K ) A (s1 ) 1 m s (MPa) n
63Sn37Pb
62Sn36Pb2Ag
Sn3.5Ag0.75Cu
Sn2.0Ag0.5Cu
9400 1:09 107 0.07 0.316 0.998 1:37 105
9400 8:49 106 0.065 0.322 0.990 6:71 104
8400 4:61 106 0.038 0.162 1.04 4:60 103
8400 2:42 107 0.043 0.168 1.005 8:10 104
Table 2 Parameters of strain hardening factor for 63Sn37Pb, 62Sn36Pb2Ag, Sn3.5Ag0.75Cu and Sn1.0Ag0.5Cu solder Parameter
63Sn37Pb
62Sn36n2Ag
Sn3.5Ag0.75Cu
Sn1.0Ag0.5Cu
a h0 (MPa)
1.26 2884
1.33 2951
1.56 3090
1.59 3162
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Fig. 9. Configuration of the test specimen.
Fig. 7. Strain rate behavior of Sn3.5Ag0.75Cu solder.
Fig. 8. Strain rate behavior of Sn1.0Ag0.5Cu solder.
affects the mechanical integrity of solder joints. Therefore it is important to study the effect of IMCs on the mechanical properties of the solder joints. The specimen consists of two copper plates with 0.9 mm thickness, plating layers and solder. The configuration of the test specimen is depicted in Fig. 9. The plating layers were composed of Cu or Ni/Cu or Au/Ni/ Cu structures. Each layer was grown by electroplating.
Cu and Ni layers were 5 lm and Au layer was less than 1 lm thickness, respectively. In this work, Sn–Ag-based solders such as Sn3.5Ag0.75Cu, Sn1.0Ag0.5Cu were used. To compare with conventional solders, Pb–Snbased solders such as 63Sn37Pb, 62Sn36Pb2Ag were also used. Strips of these solders were put between copper plates. Then, infra-red reflow process which is used in standard surface mount technology (SMT) was carried out for the reflow. The reflow temperatures are shown in Table 3. To study the solid states growth kinetics of the IMCs at the solder joints, specimens were placed in the isothermal chamber at 125 °C for 1–113 days. After cooling down for one day at room temperature, tensile tests were performed. The tensile load was applied perpendicular to the solder joint layer with cross-head speed of 0.5 mm/min at 22 °C. After tensile tests, fracture surface was observed by using the SEM. For microstructural observation at the solder joint interface, some of the specimens were sectioned using a low-speed diamond saw and metallographically polished to reveal the interface and the internal microstructure of the solder joints. Microstructural characterization was curried out and the composition of each phase was investigated by EDX microprobe analysis. 3.1. Growth of intermetallic compound layer Cross-sectional morphologies of the intermetallic layers in Cu/62Sn36Pb2Ag solder joints annealed for 0 and 80 days at 125 °C are shown in Fig. 10. Back-scattered
Table 3 Solder types and reflow temperatures Solder types
63Sn37Pb
62Sn26Pb2Ag
Sn3.5Ag0.75Cu
Sn1.0Ag0.5Cu
Melting temperature
183
Reflow temperature Plating metals
240 Cu, Ni, Ni þ Au
170 190 240
216 220 260
219 216 260
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Fig. 10. Microstructures of Cu/62Sn36Pb2Ag solder interface.
electron imaging (BEI) was used to obtain the micrographs, in order to produce better contrast among various layers of material. From left to right, the regions in each of these micrographs are the solidified solder, reaction zone, and Cu layer. IMC region can be seen clearly in the 80 days aged specimen. From EDX microprobe analysis, g-phase Cu6 Sn5 and e-phase Cu3 Sn were found in the reaction zone. Such microstructures of Cu–Sn interfacial IMCs described above are similar to what other researchers have found in bimetallic couples of Sn–Pb solders on copper [19]. The formation of g-phase Cu6 Sn5 intermetallic layers in solder joint during the reflow process arises by interfacial reactions between its constituting species, Sn from the solder and Cu from the copper plate. During isothermal aging, the g-phase Cu6 Sn5 in IMC layer grows by interdiffusion of Cu and Sn and reaction with each other, while the g-phase Cu3 Sn forms and grow by reactions between the Cu substrate and g-phase Cu6 Sn5 layer.
Furthermore, Pb-rich phases were educed adjacent to the intermetallic compounds due to the migration of Sn from solder to the reaction zone. Ni metallization is commonly used as a protective layer on a Cu conductor in electronic devises and circuit fabrications. This layer provides a diffusion barrier to inhibit detrimental growth of Cu–Sn IMCs. Interfacial reaction between 62Sn36Pb2Ag and Ni plating layer is depicted in Fig. 11. Ni/62Sn36Pb2Ag solder joints were aged for 0 and 80 days at 125 °C. From left to right, the regions in each of these micrographs are the solidified solder, reaction zone, Ni layer and Cu layer. IMCs, which were formed between Ni plating and the solder, can be seen clearly in the 80 days aging specimen. The Pb-rich phases were educed adjacent to the IMC layer like Cu plating case. The thickness of IMCs is thinner than that of Cu plating case. From EDX microprobe analysis, IMC layer was composed of Ni–Sn compounds.
Fig. 11. Microstructures of Ni/62Sn36Pb2Ag solder interface.
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Fig. 12. Microstructures of Au–Ni/62Sn36Pb2Ag solder interface.
The Au–Ni–Cu three layer structure is one of the most common solder ball pad finishes for the ball-gridarray (BGA) packages. Interfacial reaction between 62Sn36Pb2Ag and Au–Ni plating layer is shown in Fig. 12. Au þ Ni/62Sn36Pb2Ag solder joints were aged for 0 and 80 days at 125 °C. From left to right, the regions in each of these micrographs are the solidified solder, reaction zone, Ni layer and Cu layer. The Au layer could not be identified here. IMCs region can be seen clearly in the 80 days aged specimen. The Pb-rich phases were educed adjacent to the IMC layer like the previous cases. The thickness of IMCs is almost equal to Ni plating case. From EDX microprobe analysis, IMC layer was composed of Ni–Sn compounds. In Au–Ni plating cases, the Au layer could not be identified from the SEM observation, and also could not be detected by EDX analysis. From the recent study of Kim and Tu [20], Au reacts rapidly with molten Pb–Sn solders to form AuSn4 . This reaction was extremely fast. At 200 °C, the AuSn4 layer could grow to
10 lm thick in 5 s. During the fabrication of the specimen, the Au layer dissolved into the solders very quickly and the Ni layer was exposed. The exposed Ni reacted with the solder to form Ni–Sn compounds at the interface. Therefore, the composition of IMC layer is similar to that of Ni plating case. Figs. 13–15 show the cross-sectional morphologies of Sn3.5Ag0.75Cu solder joints treated by each metallization. The aging periods are for 0 and 80 days at 125 °C. IMCs region can be seen clearly in the 80 days aged specimen. Since this solder is Pb-free type, Pb-rich phases could not be seen at the interface. The thickness of IMC layer is thinner than that of 62Sn36Pb2Ag solder joints for every plating cases. The composition of IMC layer is similar to 62Sn36Pb2Ag solder joints. The thickness of IMC layers for 63Sn37Pb solder and Sn1.0Ag0.5Cu solder with Cu or Ni plating case is depicted in Fig. 16. From this figure, the thickness of IMC layers increases in proportion to the square of aging period. The thickness of IMC layers of Pb-contained
Fig. 13. Microstructures of Cu/Sn3.5Ag0.75Cu solder interface.
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Fig. 14. Microstructures of Ni/Sn3.5Ag0.75Cu solder interface.
Fig. 15. Microstructures of Au þ Ni/Sn3.5Ag0.75Cu solder interface.
solder is larger than that of Pb-free solder. The reaction rate of Ni and Sn is slower than that of Cu and Sn. For the Pb-free solder joints with Ni plating, the IMCs thickness is almost constant during this aging period. IMC layer was formed at reflow process stage, which it is not influenced by the aging periods under 125 °C. This result suggests that, for Pb-free solder, the IMC development can be controlled by using Ni plating pads. 3.2. Tensile strength of the aged solder joints
Fig. 16. Development of the IMC layers.
From the results of the cross-sectional morphologies, several kinds of IMC layers were formed at the interface. The formation of these compounds seems to affect the reliability of the solder joints. In this study, tensile tests were performed for the solder joints after thermal aging and the relation between IMCs growth and the tensile strength were investigated.
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From the observation of the tested specimen, the fracture pattern can be classified under the three types as shown in Fig. 17. Fig. 17(a) shows that the fracture occurs between solder and IMC layer and Fig. 17(b) shows the fracture occurred in the solder itself. Mixed fracture type of these two types is shown in Fig. 17(c). Fig. 18 shows the relation between tensile strength and thermal aging time for Cu plating case. The symbols used in this figure represent the fracture patterns classified in Fig. 17. From this figure, the tensile strengths of the Pb-contained solders are larger than that of the Pb-free solders. As increasing Ag weight (%) in the solder, tensile strength tends to increase. This is the same trend as the results of tensile testing for solders themselves. The tensile strengths decrease with aging period and they become almost constant if the aging period is over 50 days. The decrease rate of the tensile strength for the Pb-free solders is larger than that of Pb-contained solders. For the Pb-contained solders, fracture occurred in the solder at first and after some aging period, fracture pattern changed to the interfacial fracture. On the other hand, fracture pattern of the Pb-free solders was mainly the interfacial fracture.
Fig. 17. Fracture patterns of failed specimen of tensile testing.
959
Fig. 18. Effect of aging time on tensile strength for Cu plating case.
The IMC layer development strongly influences the tensile strength of the Pb-free solder joints. Therefore, Cu plating pad finishes are not preferable for the Pb-free solders. Fig. 19 shows the relation between the tensile strength and thermal aging time for Ni plating case. During short aging period, the tensile strengths of the Pb-contained solders are larger than that of the Pb-free solders. However, as the aging time increase, the strengths of the Pb-contained solders decrease and they become almost constant if the aging time period is over 50 days. For the Pb-free solders, the tensile strengths were almost constant during this aging period. After long time aging, the strengths of the Pb-free solder become larger than
Fig. 19. Effect of aging time on tensile strength for Ni plating case.
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Fig. 20. Effect of aging time on tensile strength for Au þ Ni plating case.
that of the Pb-contained solder. For the Pb-contained solders, fracture occurred in the solder at first and after some aging period, fracture pattern tend to be the interfacial fracture. On the other hand, for the Pb-free solders, the fracture occurred in the solders themselves. The interface strengths between solder and Ni are larger than the strengths of the solders themselves. These trends are different from Cu plating case. The effect of aging time on the tensile strength for Au and Ni plating case is depicted in Fig. 20. During short aging period, the tensile strengths of all systems are almost same. As the aging time increase, the strengths of the Pb-contained solders decrease and become constant after 50 days aging. On the other hand, the strengths of the Pb-free solders are almost constant. After short time aging (less than one month), the strengths of the Pb-free solder become larger than those of the Pb-contained
solders. Fracture patterns of the Pb-free solders are mainly the interfacial fracture. The interface strengths between solder and Ni are smaller than the strengths of the solders themselves. These trends are different from Ni plating case. From the SEM and EDX analysis, the IMCs of Au/Ni plating case were similar to Ni plating case. As mentioned before, Au diffused into the solder during reflow process. Ho et al. [21] investigated this phenomenon in detail. At first stage in reflow process, the Au layer reacted very quickly with the solder to form AuSn4 and AuSn2 . AuSn2 is immediately converted to AuSn4 . The growth ratio of AuSn4 and AuSn2 was nearly 1 lm/s. Then, AuSn4 grains began to separate themselves from the Ni layer at the roots of the grains and started to fall into the solder. After AuSn4 left the interface, the Ni layer reacted with Sn to form Ni–Sn compounds. AuSn4 was dispersed into the solder and this distributed AuSn4 contributes to the enhancement of the tensile strength of the solders. Therefore, the interface strengths between solder and Ni become smaller than the strengths of the solders and fracture pattern changes to the interfacial fracture. The reaction rate of Ni and Sn is very slow and the surface condition at the interface is not so different during this aging period. Therefore, thermal aging do not affect the tensile strength of Pb-free solders with Au and Ni pads finishes and this pads finishes is preferable for Pb-free solders. 3.3. Morphology of the fracture surface Fig. 21 shows the fractography of the Cu/62Sn36Pb2Ag solder joints. In this case, fracture occurred at the interface between solder and IMC layer. Before aging, many small dimples can be observed and ductile rupture occurred at the interface. After thermal aging, these dimples become large and the fracture
Fig. 21. Fracture surface of Cu/62Sn36Pb2Ag solder joints.
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surface was filled with these dimples. This is due to the formation of the g-phase Cu6 Sn5 [22]. At first, columnar grains of Cu6 Sn5 are formed on the Cu plates and as the aging times increase, the growth of these columnar grains proceeds through coarsening of the grains. As the dimples become large, the tensile strength decreases and when they become large enough, the tensile strength becomes constant. Fig. 22 shows the fractography of the Ni/62Sn36Pb2Ag solder joints. Before aging, fracture occurred in the solder. The fracture surface is flat and brittle fracture seems to occur due to the plastic constraint. After aging, the fracture type changes to the interfacial fracture. The needle-like Ni3 Sn4 can be observed in the fracture surface. The fracture is supposed to occur inside the IMC layer. Therefore, once the IMC layer formed, the tensile strength depends on the strength of IMC layer itself. The fracture surfaces of the Au þ Ni/62Sn36Pb2Ag solder joints are shown in Fig. 23. In this case, fracture occurred at the interface. Before
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aging, several dimples can be observed and ductile fracture is supposed to occur in this case. After aging, the fracture surface becomes flat and the brittle fracture could occur. The fracture mode is considered to be changed due to thermal aging. Therefore, the tensile strength of this case decreases with aging period. Fig. 24 shows the fractography of Cu/Sn3.5Ag0.75Cu solder joints. In this case, the fracture occurred at the interface. There are many dimples and those dimples become large after thermal aging. These dimples are due to the formation of the Cu6 Sn5 like 62Sn36Pb2Ag case. The fracture surfaces of the Ni/62Sn36Pb2Ag solder joints are shown in Fig. 25. In these cases, the fracture occurred inside the solder before and after aging. The fracture surface is flat and similar to each other. Brittle fracture is supposed to occur in these cases and the tensile strength does not depend on the aging period. Fig. 26 shows the fractography of Au þ Ni/62Sn36Pb2Ag solder joints. In these cases, the fracture occurred at the interface. The
Fig. 22. Fracture surface of Ni/62Sn36Pb2Ag solder joints.
Fig. 23. Fracture surface of Au þ Ni/62Sn36Pb2Ag solder joints.
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Fig. 24. Fracture surface of Cu/Sn3.5Ag0.75Cu solder joints.
Fig. 25. Fracture surface of Ni/Sn3.5Ag0.75Cu solder joints.
Fig. 26. Fracture surface of Au þ Ni/Sn3.5Ag0.75Cu solder joints.
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fracture surface is flat and similar to each other. Brittle fracture is supposed to occur in these cases and the tensile strength does not depend on the aging period.
3.4. Board level reliability tests To verify the variation of specimen samples used in the aging tests at 125 °C, several studies have been conducted using chip scale packages (CSP)-mounted FR-4 board. In this study, the CSP assembles with polyimide-based substrates were tested under thermal cyclic loads. The 10 10 mm2 CSP consists of the silicon chip (6:0 6:0 0:28 mm3 ), an epoxy-based molding compound and 151 solder balls at a 0.5 mm pitch. Two types of solder ball materials were used, 62Sn36Pb2Ag and Sn1.0Ag0.5Cu, respectively. 40 to 125 °C thermal cycles with 5 min ramp and 10 min dwell times were used. The 0.8 mm thick PCB having the Au/Ni or copper pad finish was used. Two types of solder pastes for the PCB pad finish were used, 63Sn37Pb and Sn3.5Ag0.75Cu, respectively. Figs. 27–30 show the three-parameter (3-P) Weibull distribution for the comparison of 62Sn36Pb2Ag versus Sn1.0Ag0.5Cu using the Au–Ni pad finish and 63Sn37Pb solder paste, the Au–Ni pad finish and Sn3.5Ag0.75Cu solder paste, the copper pad finish and 63Sn37Pb solder paste, the copper pad finish and 63Sn37Pb solder paste, respectively. The 3-P Weibull distribution is given by [23]. F ðxÞ ¼ 1 exp
xa Nf ð63:2%Þ a
Fig. 28. Comparison of 62Sn36Pb2Ag versus Sn1.0Ag0.5Cu using the Au–Ni pad finish and Sn3.5Ag0.75Cu solder paste.
Fig. 29. Comparison of 62Sn36Pb2Ag versus Sn1.0Ag0.5Cu using the copper pad finish and 63Sn37Pb solder paste.
b ! ð11Þ
where x is the number of cycles, a is the failure-free life, Nf (63.2%) is the characteristic life and b is the shape parameter which indicates the amount of scatter in the data. The parameters of 3-P Weibull, a; b and 63.2%, are obtained by non-linear least-squares fitting method.
Fig. 30. Comparison of 62Sn36Pb2Ag versus Sn1.0Ag0.5Cu using the copper pad finish and Sn3.5Ag0.75Cu solder paste.
Fig. 27. Comparison of 62Sn36Pb2Ag versus Sn1.0Ag0.5Cu using the Au–Ni pad finish and 63Sn37Pb solder paste.
a is determined by the value of R2 which is a non-linear square fitting parameter closing to 1.0 as much as possible. For the Au–Ni pad finish, Sn1.0Ag0.5Cu solder balls have a factor of 1.16x and 1.52x solder fatigue life (0.1% cumulative fails) of 62Sn36Pb2Ag balls using 63Sn37Pb and Sn3.5Ag0.75Cu solder paste, respectively. On the
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contrary, for the copper pad finish, Sn1.0Ag0.5Cu solder balls have a factor of 0.89xand 0.80x solder fatigue life (0.1% cumulative fails) of 62Sn36Pb2Ag balls using 63Sn37Pb and Sn3.5Ag0.75Cu solder paste, respectively. Based on the results of board level reliability tests, it is found that the Au–Ni pad finish makes Sn1.0Ag0.5Cu a more solder fatigue life than 62Sn36Pb2Ag while the copper pad finish decreases the life of Sn1.0Ag0.5Cu against 62Sn36Pb2Ag. The results of board level reliability tests are similar phenomenon to the results of simple specimen samples used in the aging tests at 125 °C. EPMA was used to determine the density of gold around the interface between the Sn1.0Ag0.5Cu solder and the Ni–Au pad finish. Figs. 31a and b show the density of gold around the interface after the soldering process (a) and few thermal cycles, respectively. After the soldering process, the gold was diffused into the solder. Following few thermal cycles, the gold was reached to the interface once again. It is expected that this phenomenon is due to the following assumption: (1) High stresses are generated
Fig. 32. Strength of 63Sn37Pb as a function of gold weight (%).
near the interface between the solder and the solder pad after cooling down to room temperature and then few thermal cycles, leading to atomic level and solder grain level defects due to dislocation caused by solder creep. (2) The gold, which is diffused into the solder, is moved to the interface once again to repair the atomic level and solder grain level defects. In board level reliability tests, fractures in solder were observed for Sn1.0Ag0.5Cu with the copper pad finish while fractures near the interface were observed for the solder with the Au–Ni pad finish. Fig. 32 reveals that the strength of 63Sn37Pb as a function of gold weight (%) under tensile and shear loads [24,25]. It is obvious that the tensile strength of 63Sn37Pb is increased in range from 0 to 4 wt% while the shear strength is increased in range from 0 to 2 wt%. Without the gold, cracks in the solder are generated for Sn1.0Ag0.5Cu. After the gold diffusion into the solder, due to an increase in the strength of Sn1.0Ag0.5Cu like 63Sn37Pb, and cracks appeared near the interface. For 62Sn36Pb2Ag, fractures near the interface were observed for both the Au/ Ni and copper pad finish. It is found that cracks in the solder are never generated because of high strength (based on the results of the tensile tests in the ratedependent plasticity study). Since cracks near the interface are generated for 62Sn36Pb2Ag, solder fatigue life was determined by microstructural change near the interface induced by the component of solder with a plating material.
4. Conclusion Fig. 31. (a) Density of gold around the interface after the soldering process. (b) Density of gold around the interface after the 125 °C aging process (white color is gold).
A rate-dependent plasticity and the strength of solders with a variety of plating materials such as Cu, Ni,
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and Ni–Au platings were characterized for Sn–Ag-based lead-free solders. In the rate-dependent plasticity characterization, the material parameters of a viscoplastic constitutive model for Sn3.5Ag0.75Cu and Sn1.0Ag0.5Cu solder alloy were determined from experimental results. This constitutive model unifies both rate-dependent creep behavior and rate-independent plastic behavior occurring concurrently at the same time in the material. A good agreement between the experimental data and the model predictions is achieved for the temperature of the strain rate behavior ranging from 25 to 125 °C and the strain rate ranges from 1 to 100 mm/min. In the study of the strength of solders with a variety of plating materials, tensile tests were performed for the solder joints after thermal aging and the relations between solder joint strength and aging periods were examined. From SEM microscope observation and EDX microprobe analysis, the growth and components of the IMC layer were also examined. For the Pb-contained solders, Cu plating is preferable of the Cu pads finishes, since the strength of tensile strength of Cu plating is larger than the other platings and is not influenced by thermal aging. On the contrary, for the Pb-free solders, Au/Ni plating is preferable of the Cu pads finishes, since the tensile strength of Au/Ni plating is larger than the other plating and is not influenced by thermal aging. In board level reliability tests using 62Sn36Pb2Ag and Sn1.0Ag0.5Cu, the variation of specimen samples used in the aging tests at 125 °C was verified. References [1] Lau John H, Pao Yi-Hsin. Solder joint reliability of BGA, CSP, Flip Chip, and Fine Pitch SMT assemblies. McGrawHill, 1997. p. 1–405. [2] Amagai M. The effect of polymer die attach material on solder joint reliability. In: Proceedings of the ASME Workshop on Mechanical Reliability of Polymeric Materials of IC Devices, Paris, France, November 1998. p. 223–30. [3] Amagai M. Chip scale package solder joint reliability. In: Proceedings of the IEEE 36th International Reliability Physics Symposium, Reno, USA, April 1998. p. 260–8. [4] Darveaux R. Solder joint fatigue life model. In: Design and reliability of solders and solder interconnections. The Minerals, Metals and Materials Society (TMS); 1997. p. 213–8. [5] Amagai M. Chip scale package (CSP) solder joint reliability and modeling. Microelectron Reliab 1999;39:463–77. [6] Amagai M. Chip scale package (CSP) material characterization. Microelectron Reliab 1999;42:573–87. [7] Amagai M. Characterization of molding compound and die attach materials for package warpage and solder joint reliability in chip scale package. In: Proceedings of the ASME InterPack, Hawaii, USA, June 1999. p. 1103–12.
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