The Effects of Aging Temperature on SAC Solder Joint ... - IEEE Xplore

The Effects of Aging Temperature on SAC Solder Joint Material Behavior and Reliability Yifei Zhang, Zijie Cai, Jeffrey C. Suhling, Pradeep Lall, Michael J. Bozack Center for Advanced Vehicle Electronics Auburn University Auburn, AL 36849 Phone: +1-334-844-3332 FAX: +1-334-844-3307 E-Mail: [email protected] that the creep behaviors of lead free and tin-lead solders experience a “cross-over point” where the lead free solders begin to creep at higher rates than standard 63Sn-37Pb solder for the same stress level. These cross-over points were observed to exist for all four of the SAC solder alloys, and for all of the aging temperatures except for room temperature. In addition, the lower silver content alloys (e.g. SAC105) were observed to be much more sensitive to aging (have greater changes in the creep rate for a given aging time) than the higher silver content alloys (e.g. SAC405). The times required before the cross-overs occurred were reduced when considering higher aging temperatures or SAC alloys with lower silver content. The cross-over points are definitely present for all of the SAC alloys when they are subjected to 100 C aging present in typical computer server environments.

Abstract The microstructure, mechanical response, and failure behavior of lead free solder joints in electronic assemblies are constantly evolving when exposed to isothermal aging and/or thermal cycling environments. In our prior work on aging effects (Ma, et al., ECTC 2006 and ECTC 2007), we demonstrated that the observed material behavior variations of SAC305 and SAC405 lead free solders during room temperature aging (25 oC) and elevated temperature aging (125 oC) were unexpectedly large and universally detrimental to reliability. Such effects for lead free solder materials are especially important for the harsh applications environments present in high performance computing and in automotive, aerospace, and defense applications. However, there has been little work in the literature, and the work that has been done has concentrated on the degradation of solder ball shear strength (e.g. Dage Shear Tester). Current finite element models for solder joint reliability during thermal cycling accelerated life testing are based on traditional solder constitutive and failure models that do not evolve with material aging. Thus, there will be significant errors in the calculations with the new lead free SAC alloys that illustrate dramatic aging phenomena. In the current work, we have extended our previous studies to include a full test matrix of aging temperatures and solder alloys. The effects of aging on mechanical behavior are being examined by performing stress-strain and creep tests on four different SAC alloys (SAC105, SAC205, SAC305, SAC405) that were aged for various durations (0-12 months) at room temperature (25 oC), and several elevated temperatures (50, 75, 100, and 125 oC). Analogous tests are being performed with 63Sn-37Pb eutectic solder samples for comparison purposes. Variations of the mechanical and creep properties (elastic modulus, yield stress, ultimate strength, creep compliance, etc.) are being observed and modeled as a function of aging time and aging temperature. In addition, the chosen selection of SAC alloys has allowed us to explore the effects of silver content on aging behavior (we have examined SACN05 with N= 1%, 2%, 3%, and 4% silver; with all alloys containing 0.5% copper). In this paper, we concentrate on presentation of the creep results from our ongoing studies. The results obtained in this work have demonstrated the significant effects of elevated temperature exposure on the creep behavior of solder joints. As expected, the creep rates evolved more dramatically when the aging temperature was increased. In addition, the effects of aging were shown to be significant even for aging temperature slightly above room temperature (e.g. T = 75 oC). The recorded data demonstrate

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Introduction Eutectic or near eutectic tin/lead (Sn/Pb) solder (melting temperature TM = 183 °C) has been the predominant choice of the electronics industry for decades due to its outstanding solderability and reliability. However, legislation that mandates the banning of lead in electronics has been actively pursued worldwide during the last 15 years due to the environmental and health concerns. Although the implementation deadlines and products covered by such legislation continue to evolve, it is clear that laws requiring conversion to lead-free electronics are becoming a reality. Other factors that are affecting the push towards the elimination of lead in electronics are the market differentiation and advantage being realized by companies producing so-called “green” products that are lead-free. A large number of research studies are currently underway in the lead-free solder area. Although no “drop in” replacement has been identified for all applications; Sn-Ag, Sn-Ag-Cu (SAC), and other alloys involving elements such as Sn, Ag, Cu, Bi, In, and Zn have been identified as promising replacements for standard 63Sn-37Pb eutectic solder. Several SAC alloys have been the proposed by various user groups and industry experts. These include 96.5Sn-3.0Ag-0.5Cu (SAC305), 95.5Sn-3.8Ag-0.7Cu (SAC387), 95.5Sn-3.9Ag0.6Cu (SAC396) and 95.5Sn-4.0Ag-0.5Cu (SAC405). For enhanced reliability during high strain rate exposures (e.g. shock and drop), several alloys with lower silver content have been recommended including 98.5Sn-1.0Ag-0.5Cu (SAC105) and 99Sn-0.3Ag-0.7Cu (SAC0307). The main benefits of the various SAC alloy systems are their relatively low melting temperatures compared with the 96.5Sn-3.5Ag binary eutectic alloy, as well as their superior mechanical and solderability properties when compared to other lead free solders.

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2008 Electronic Components and Technology Conference

Solder joint fatigue is one of the predominant failure mechanisms in electronic assemblies exposed to thermal cycling. Reliable, consistent, and comprehensive solder constitutive equations and material properties are needed for use in mechanical design, reliability assessment, and process optimization. Mechanical characterization of solder materials has always been hampered by the difficulties in preparing test specimens that reflect the same true material making up the as actual solder joints (e.g. match the solder microstructure). Solder uniaxial samples haven been fabricated by machining of bulk solder material [1-8], or by melting of solder paste in a mold [9-18]. Use of a bulk solder bars is undesirable, because they will have significantly different microstructures than those present in the small solder joints used in microelectronics assembly. In addition, machining can develop internal/residual stresses in the specimen, and heat generated during turning operations can cause localized microstructural changes on the exterior of the specimens. Reflow of solder paste in a mold causes challenges with flux removal, minimization of voids, microstructure control, and extraction of the sample from the mold. In addition, many of the developed specimens have shapes that significantly deviate from being long slender rods. Thus, undesired nonuniaxial stress states will be produced during loading. Other investigators have attempted to extract constitutive properties of solders by direct shear or tensile loading [6, 1928], or indenting [29-32], of actual solder joints (e.g. flip chip solder bumps or BGA solder balls). While such approaches are attractive because the true solder microstructure is involved, the unavoidable non-uniform stress and strain states in the joint make the extraction of the correct mechanical properties or stress-strain curves from the recorded loaddisplacement data very challenging. Also it can be difficult to separate the various contributions to the observed behavior from the solder material and other materials in the assembly (bond pads, silicon die, PCB/substrate, etc.). The microstructure, mechanical response, and failure behavior of lead free solder joints in electronic assemblies are constantly evolving when exposed to isothermal aging and/or thermal cycling environments [8, 13, 15-16, 18, 23-24, 28, 31, 33-49]. The observed material behavior variation during thermal aging/cycling is universally detrimental to reliability and includes reductions in stiffness, yield stress, ultimate strength, and strain to failure, as well as highly accelerated creep. Such aging effects are greatly exacerbated at higher temperatures typical of thermal cycling qualification tests. However, significant changes occur even with aging at room temperature [13, 15-16, 23-24, 33-41, 48]. As early as 1956, Medvedev [33] observed a 30% loss of tensile strength for bulk solder Sn/Pb solder stored for 450 days at room temperature. In addition, he reported 4-23% loss of tensile strength for solder joints subjected to room temperature storage for 280-435 days. In 1976, Lampe [34] found losses in shear strength and hardness of up to 20% in Sn-Pb and Sn-Pb-Sb solder alloys stored for 30 days at room temperature. Miyazawa and Ariga [35-36] measured significant hardness losses and microstructural coarsening for Sn-Pb, Sn-Ag, and Sn-Zn eutectic solders stored at 25 °C for 1000 hours, while Chilton and co-workers [37] observed a 1015% decrease in fatigue life of single SMD joints after room

temperature aging. Several studies [38-41] have also documented the degradation of Sn-Pb and SAC solder ball shear strength (10-35%) in area array packages subjected to room temperature aging. The effects of room temperature isothermal aging on constitutive behavior have also been investigated [13, 15-16, 48]. Chuang, et al. [13] characterized the reductions in yield stress and increases in elongations obtained in Sn-Zn eutectic solder during aging at room temperature. In addition, Xiao and Armstrong [15-16] recorded stress-strain curves for SAC 396 specimens subjected to various durations of room temperature aging, and found losses of ultimate tensile strength of up to 25%. The effects of room temperature aging on the mechanical properties and creep behavior of SAC alloys have been extensively discussed by the authors (Ma, et. al. [48]). The measured stress-strain data demonstrated large reductions in stiffness, yield stress, ultimate strength, and strain to failure (up to 40%) during the first 6 months after reflow solidification. In addition, even more dramatic evolution was observed in the creep response of aged solders, where up to 100X increases were found in the steady state (secondary) creep strain rate (creep compliance) of lead free solders that were simply allowed to sit in a room temperature environment. The SAC solder materials in room temperature aged joints were also found to enter the tertiary creep range (imminent failure) at much lower strain levels than virgin joints (non aged, immediately after reflow solidification). We also demonstrated that there are corresponding changes in the solder joint microstructure occurring during room temperature aging. The magnitudes of the material behavior evolution occurring in lead free SAC solder joints were found to be much larger (e.g. 25X) than the corresponding changes occurring in traditional Sn-Pb assemblies. The effects of aging at elevated temperature are the most widely studied due to the dramatic changes in the microstructure and mechanical properties that result. Aging effects (reduced effective stiffness and ultimate strength) have been observed for solder subjected to elevated temperature aging (e.g. 125 oC) [8, 15-16, 18, 23-24]. Pang, et al. [20] measured microstructure changes, intermetallic layer growth, and shear strength degradation in SAC single ball joints subjected to elevated temperature aging. Darveaux [21] performed an extensive experimental study on the stress-strain and creep behavior of area array solder balls subjected to shear. He found that aging for 1 day at 125 oC caused significant effects on the observed stress-strain and creep behavior. The aged specimens were also found to creep much faster than un-aged ones by a factor of up to 20 times for both SAC305 and SAC405 solder alloys. Xiao and Armstrong [1314] measured stress-strain curves for SAC396 specimens subjected to elevated temperature aging at 180 oC. At this highly elevated temperature, they observed a quick softening of the material during the first 24 hours followed by a gradual hardening with time. Dutta, et al. [31] used impression techniques to measure the creep behavior of SAC405 solder joints and observed large increases in the secondary creep Wiese and Kolter [28] rates with aging at 180 oC. demonstrated analogous large increases in the creep rates for

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SAC387 joints by directly loading flip chip assemblies that were aged at 125 oC. Several studies have been performed on the degradation of BGA ball shear strength with elevated temperature aging at 125 oC or 150 oC [42-46]. All of these investigations documented both microstructure coarsening and intermetallic layer growth. In addition, Hasegawa, et al. [42] measured elastic modulus reductions with aging by testing thin solder wires, while Chiu and co-workers [46] found significant reductions in drop reliability during elevated temperature aging. Finally, Ding, et al. [47] explored the evolution of fracture behavior of SnPb tensile samples with elevated temperature aging. In our prior work on elevated temperature aging effects (Ma, et al., [49]), we demonstrated that the observed material behavior variations of SAC305 and SAC405 lead free solders during isothermal aging at 125 oC were unexpectedly large and universally detrimental to reliability. The measured stress-strain data demonstrated large reductions in stiffness, yield stress, ultimate strength, and strain to failure (up to 50%) during the first 6 months after reflow solidification. After approximately 1000 hours of aging, the lead free solder joint material properties were observed to degrade at a slow but constant rate. In addition, even more dramatic evolution was observed in the creep response of aged solders, where up to 500X increases in the secondary creep rates were observed for aging up to 6 months. The solder materials in aged joints were also found to enter the tertiary creep range (imminent failure) at much lower strain levels than virgin joints (non aged, tested immediately after reflow solidification). We also correlated the changes in mechanical behavior during aging with changes that occur in the solder joint microstructure, and showed that the magnitudes of the material behavior evolution occurring in lead free SAC solder joints are much larger (e.g. 100X) than the corresponding changes occurring in traditional Sn-Pb assemblies. One of the most important observations from our prior work on creep behavior was the demonstration that a “crossover point” occurs during the elevated temperature aging (125 C) of lead free and tin-lead solders. This cross-over point occurred after approximately 50 hours of aging, and marked the point where the two lead free solders began to creep at higher rates than standard 63Sn-37Pb solder for the same stress level. Such an effect was not observed for solder joints aged at room temperature (25 C). The presence of the crossover point with elevated temperature aging can possibly explain existing reliability data for area array packages where lead free packaging becomes less reliable than the analogous Sn-Pb case when the upper limit of the thermal cycling test is increased. As demonstrated above, the literature has documented the dramatic changes occurring in the constitutive and failure behavior of solder materials and solder joint interfaces during isothermal aging. However, these effects have been largely ignored in most other studies involving solder material characterization or finite element predictions of solder joint reliability during thermal cycling. It is also widely acknowledged that the large discrepancies in measured solder mechanical properties from one study to another are due to differences in the microstructures of the tested samples. This

problem is exacerbated by the aging issue, as it is clear that the microstructure and material behavior of the samples used in even a single investigation are moving targets that are evolving rapidly even at room temperature. Furthermore, the effects of aging on solder behavior must be better understood so that more accurate viscoplastic constitutive equations can be developed for SnPb and SAC solders. Without such relations, it is doubtful that finite element reliability predictions can ever reach their full potential. In the current work, we have extended our previous studies to include a full test matrix of aging temperatures and solder alloys. The effects of aging on mechanical behavior are being examined by performing stress-strain and creep tests on four different SAC alloys (SAC105, SAC205, SAC305, SAC405) that were aged for various durations (0-12 months) at room temperature (25 oC), and several elevated temperatures (50, 75, 100, and 125 oC). Analogous tests are being performed with 63Sn-37Pb eutectic solder samples for comparison purposes. Variations of the mechanical and creep properties (elastic modulus, yield stress, ultimate strength, creep compliance, etc.) are being observed and modeled as a function of aging time and aging temperature. In addition, the chosen selection of SAC alloys allowed the effects of silver content on aging behavior to be explored (we have examined SACN05 with N= 1%, 2%, 3%, and 4% silver; with all alloys containing 0.5% copper). In this paper, we concentrate on presentation of the creep results from our ongoing studies. The existence of the creep cross-over point for the different alloys with aging at various temperatures above room temperature has been explored in detail. Experimental Procedure Uniaxial Test Sample Preparation In the current study, mechanical measurements of aging effects and material behavior evolution of lead free solders have been performed. We have avoided the specimen preparation pitfalls present in many previous studies by a using a novel procedure where solder uniaxial test specimens are formed in high precision rectangular cross-section glass tubes using a vacuum suction process. The tubes were then cooled by water quenching and sent through a SMT reflow to re-melt the solder in the tubes and subject them to any desired temperature profile (i.e. same as actual solder joints). The solder is first melted in a quartz crucible using a pair of circular heating elements (see Figure 1). A thermocouple attached on the crucible and a temperature control module is used to direct the melting process. One end of the glass tube is inserted into the molten solder, and suction is applied to the other end via a rubber tube connected to the house vacuum system. The suction forces are controlled through a regulator on the vacuum line so that only a desired amount of solder is drawn into the tube. The specimens are then cooled to room temperature using a user-selected cooling profile. In order to see the extreme variations possible in the mechanical behavior and microstructure, we are exploring a large spectrum of cooling rates including water quenching of the tubes (fast cooling rate), air cooling with natural and forced convection (slow cooling rates), and controlled cooling using a surface mount technology solder reflow oven. Typical temperature versus time plots for water quenching and air

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cooling of the test samples are shown in Figure 2. For the reflow oven controlled cooling, the tubes are first cooled by water quenching, and they are then sent through a reflow oven (9 zone Heller 1800EXL) to re-melt the solder and subject it to the desired temperature profile. Thermocouples are attached to the glass tubes and monitored continuously using a radio-frequency KIC temperature profiling system to ensure that the samples are formed using the desired temperature profile (same as actual solder joints). Figure 3 illustrates the reflow temperature profiles used in this work for SAC and SnPb solder specimens.

can be filled with solder, and the cross-sectional dimensions of the hole running the length of the tube. In the current work, we formed uniaxial samples with nominal dimensions of 80 x 3 x 0.5 mm. A thickness of 0.5 mm was chosen because it matches the height of typical BGA solder balls.

(a) SAC (105/205/305/405)

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Figure 1 - Specimen Preparation Hardware

Figure 3 - Solder Reflow Temperature Profiles

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(b) Air Cooled Figure 2 - Sample Cooling Profiles Typical glass tube assemblies filled with solder and a final extracted specimen are shown in Figure 4. For some cooling rates and solder alloys, the final solidified solder samples can be easily pulled from the tubes due to the differential expansions that occur when cooling the low CTE glass tube and higher CTE solder alloy. Other options for more destructive sample removal involve breaking the glass or chemical etching of the glass. The final test specimen dimensions are governed by the useable length of the tube that

The described sample preparation procedure yielded repeatable samples with controlled cooling profile (i.e. microstructure), oxide free surface, and uniform dimensions. By extensively cross-sectioning several specimens, we have verified that the microstructure of any given sample is consistent throughout the volume of the sample. In addition, we have established that our method of specimen preparation yields repeatable sample microstructures for a given solidification temperature profile. Samples were inspected using a micro-focus x-ray system to detect flaws (e.g. notches and external indentations) and/or internal voids (non-visible). Figure 5 illustrates results for good and poor specimens. With

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proper experimental techniques, samples with no flaws and voids were generated.

Figure 5 - X-Ray Inspection of Solder Test Specimens (Good and Bad Samples) Mechanical Testing System A MT-200 tension/torsion thermo-mechanical test system from Wisdom Technology, Inc., as shown in Figure 6, has been used to test the samples in this study. The system provides an axial displacement resolution of 0.1 micron and a rotation resolution of 0.001°. Testing can be performed in tension, shear, torsion, bending, and in combinations of these loadings, on small specimens such as thin films, solder joints, gold wire, fibers, etc. Cyclic (fatigue) testing can also be performed at frequencies up to 5 Hz. In addition, a universal 6-axis load cell was utilized to simultaneously monitor three forces and three moments/torques during sample mounting and testing. Environmental chambers added to the system allow samples to be tested over a temperature range of approximately -185 to +300 °C.

the specimen length to width aspect ratio was 20 to 1 (insuring true uniaxial stress states). All uniaxial stress-strain and creep tests in this paper were conducted at room temperature (25 °C). The strain rate for the stress-strain testing was ε& = 0.001 sec −1 , while the applied stress for the creep testing was σ = 15 MPa. Typical Test Data A typical recorded tensile stress strain curve for solder with labeled standard material properties is shown in Figure 7. The notation “E” is taken to be the effective elastic modulus, which is the initial slope of the stress-strain curve. Since solder is viscoplastic, this effective modulus will be rate dependent, and will approach the true elastic modulus as the testing strain rate approaches infinity. The yield stress σ Y (YS) is taken to be the standard .2% yield stress (upon unloading, the permanent strain is equal to ε = .002 ). Finally, the ultimate tensile strength σ u (UTS) is taken to be the maximum stress realized in the stress-strain data. As shown in Figure 7, the solders tested in this work illustrated nearly perfect elastic-plastic behavior (with the exception of a small transition region connecting the elastic and plastic regions). As the strain level becomes extremely high and failure is imminent, extensive localized necking takes place. These visible reductions in cross-sectional area lead to non-uniform stress-states in the specimen and drops in the applied loading near the end of the stress-strain curve. σu Stress

σY

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Strain

Figure 7 - Typical Solder Stress-Strain Curve and Material Properties

Figure 6 - MT-200 Testing System with Solder Sample During uniaxial testing, forces and displacements were measured. The axial stress and axial strain were calculated from the applied force and measured cross-head displacement using

σ=

F A

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∆L δ = L L

(1)

where σ is the uniaxial stress, ε is the uniaxial strain, F is the measured uniaxial force, A is the original cross-sectional area, δ is the measured crosshead displacement, and L is the specimen gage length (initial length between the grips). The gage length of the specimens in this study was 60 mm, so that

Figure 8 illustrates a typical solder creep curve (strain vs. time response for a constant applied stress). The response begins with a quick transition to the initial “elastic” strain level, followed by regions of primary, secondary, and tertiary creep. Depending on the applied stress level, the primary creep region can be more extensive for the SAC alloys relative to Sn-Pb solders. The secondary creep region is typically characterized by a very long duration of nearly constant slope. This slope is referred to as the “steady state” secondary creep rate or creep compliance, and it is often used by practicing engineers as one of the key material parameters for solder in finite element simulations used to predict solder joint reliability. In this work, the measured creep rates were taken to be the minimum slope values in the secondary creep regions of the observed ε& versus t responses. The tertiary creep region occurs when rupture is imminent, and typically features an abrupt change to a nearly constant but significantly increased creep rate.

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Figure 8 - Typical Solder Creep Curve Test Matrix Description Using specimens fabricated with the casting procedure described above, thermal aging effects and viscoplastic material behavior evolution are being characterized for SAC105, SAC205, SAC305, and SAC405 lead free solders. These alloys are commonly used for solder balls in lead free BGAs, CSPs, and flip chip die, as well as for solder pastes used in SMT processes with other components. The lead free solder test results presented in this paper are all for samples solidified with the reflowed (Figure 3a) cooling profile (mimics that seen by actual solder joints during PCB assembly). Analogous experiments are also being performed with reflowed 63Sn-37Pb eutectic solder samples for comparison purposes. In our ongoing testing program, uniaxial stress-strain curves and creep curves are being characterized for five different aging temperatures including T = 25 (RT), 50, 75, 100, and 125 °C. The test matrix of aging temperatures and aging times for each alloy is shown in Figure 9. We have also tested SAC405 with aging at T = 150 oC [50].

strength, creep compliance, etc.) with aging are observed and then modeled as a function of aging time. In the solder creep experiments in our testing program, constant stress levels on the order of 25-50% of the observed UTS are being applied. For the data presented in this paper, the applied stress was σ = 15 MPa, which is approximately 25%-33% of the non-aged UTS values for the various alloys tested. Due to the long test times involved, only 5 specimens are being tested for each alloy for any given set of aging conditions. From the recorded strain vs time curves under constant stress, the “steady state” creep strain rates have been extracted. Variations of the average creep rates with aging are observed and then modeled as a function of aging time. In this paper, we concentrate on presentation of the creep results from our ongoing studies. The test matrix of experiments presented in the paper is shown in Figure 10 (subset of Figure 9). For each of the 5 solder alloys (SAC105, SAC205, SAC305, SAC405, and Sn-Pb), we have measured the creep response after 0, 1, 2, 3, and 4 months of aging at 4 different aging temperatures (T = 25, 75, 100, and 125 oC). For each alloy and set of aging conditions, the strain versus time responses have been recorded and the creep rates in the secondary creep region have been evaluated. The evolution of the creep rate with aging time has been measured for each aging temperature, and then the results for the different alloys have been compared. Aging Temperature Aging Time (months) 0 1 2 3 4

25 C (RT)

75 C

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√ √ √ √ √

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Figure 10 - Aging Test Matrix for the Creep Experiments Presented in the Current Paper

Aging Temperature Aging Time (months) 0 0.5 1 2 3 4 5 6 12

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√ √ √ √ √ √ √ √ √

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Mechanical Testing Results

Figure 9 - Aging Test Matrix for Each Solder Alloy All uniaxial stress-strain and creep tests are being conducted at room temperature (25 °C). For the uniaxial stress-strain tests, a total of 10 specimens are being tested for each solder alloy at each set of aging conditions. From the recorded stress-strain data, a set of averaged material properties are being extracted. Variations of the average mechanical properties (elastic modulus, yield stress, ultimate

Effects of Aging on Solder Creep Response Uniaxial specimens were formed for the various SAC solder alloys using the methods described in previous sections, and then aged at T = 25, 75, 100, and 125 oC for up to four months. As mentioned above, specimens for each alloy were prepared in sets of five, which were then subjected to a specific set of aging conditions (aging temperature and aging time). Figures 11-14 illustrate the typical recorded creep curves for the SAC105, SAC205, SAC305, and SAC405 solder samples. In each figure, there are four graphs for the four different aging temperatures (T = 25, 75, 100, and 125 oC). In each plot, the various creep curves are for different aging times, illustrating the evolution of the creep response with duration of aging. For brevity and clarity of the presentation, only one of the five available creep curves is shown in each plot for each set of aging conditions.

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Figure 12 - Typical Creep Curves for SAC205 Aging Temperatures: T = 25, 75, 100, 125 oC Aging Times: t = 0, 1, 2, 3, 4 Months


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0.00

0.00 0

2000

4000

6000

8000

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2000

4000

6000

8000

Time, t (sec)

Time, t (sec)

0.08

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SAC 305, RF

SAC 405, RF

o

T = 25 oC

T = 25 C

σ = 15 MPa

σ = 15 MPa

0.06

0.06

Strain, ε

Strain, ε

8000

Time, t (sec)

Time, t (sec)

0.04


Aging Conditions As Reflowed o 125 C, 1 Month o 125 C, 2 Months o 125 C, 3 Months o 125 C, 4 Months

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Time, t (sec)

Time, t (sec)



106


similar creep-aging behavior in our previous study with a higher applied stress level of σ = 20 MPa [49].

107

SAC105, RF T = 25 oC,

10-2

σ =15 MPa

Strain Rate (sec-1)

10-3 10-4 10-5 10-6 10-7 10-8 Aging Aging Aging Aging

10-9

at RT o at 75 C o at 100 C o at 125 C

10-10 0

1

2

3

4

5

Aging Time (month)

Figure 15 - Evolution of Creep Strain Rate with Aging (SAC105; Aging at 25, 75, 100 125 oC) 10-1

SAC205, RF T = 25 oC

10-2

σ =15 MPa

-1

Strain Rate (sec )


10-9


10-10 0

1

2

3

4

5

Aging Time (month)

Figure 16 - Evolution of Creep Strain Rate with Aging (SAC205; Aging at 25, 75, 100 125 oC) 10-1


10-2

σ =15 MPa

10-3 -1

Evolution of Creep Rate with Aging Exposure The effects of aging on the creep rate can be better seen by plotting the extracted secondary creep rates versus the aging time for each alloy. Such graphs are presented for SAC105, SAC205, SAC305, and SAC405 in Figures 15-18, respectively. In each plot, the creep rate evolution is indicated for each of the four aging temperatures. Each data point represents the average creep rate measured for the 5 samples tested at a given set of aging conditions. From the results in Figures 15-18, it is apparent that all four of the SAC alloys experience dramatic changes in their creep rate for elevated temperature aging (100-5000X). As in our prior work [48-49], the creep rate data have been fit with a combination of power law and linear models to illustrate evolution of the creep rate with aging time. It is observed that the functional variations with aging at 75, 100, and 125 oC are approximately “in parallel” and are closely spaced. They are also significantly separated from the variation occurring with room temperature aging. After large changes occur during the first month of aging, the variation of the creep rate for all four SAC alloys becomes nearly linear for longer aging times. For the data we have currently (up to 4 months of aging), there is no indication that an aging saturation point will be reached where the creep rate stabilizes. This suggests that there will potentially be reliability problems for SAC solder joints subjected to long term exposures at temperatures T > 75 oC. We have additional samples currently aging for 6 and 12 months to further explore this phenomenon. A final observation from the creep rate data in Figures 1518 is that the aging effects on the creep response are much stronger for the lower silver content alloys (SAC105 and SAC205). Considering all of the tests performed so far, the maximum recorded creep rate for each alloy was obtained with 4 months aging at 125 oC (most severe aging conditions considered). Figure 19 contains numerical values of the ratio of these maximum creep rates to the corresponding creep rates for the non-aged (as reflowed material). These ratios represent the worst cases of the observed creep rate increases with aging. The changes for SAC105 and SAC205 were very large at 5500X and 1000X respectively. The changes for SAC305 and SAC405 were each 220X. Although 220X is still a very significant change, the aging effect on the creep response is significantly smaller when the SAC alloy contains more silver. We also observed that SAC305 and SAC405 had

10-1

Strain Rate (sec )

The plots in Figures 11-14 clearly indicate a dramatic evolution of the creep response of all the SAC alloys when they are aged at elevated temperatures of T = 75, 100, and 125 o C. The first plot in Figure 11 also demonstrates a significant effect on the creep response for room temperature aging of SAC105. The evolution of the creep response at room temperature is not as immediately obvious for SAC205, SAC305, and SAC405 (see first graphs in Figures 12-14). However, the evolution was indeed present although it was relatively small compared to the observed variations with elevated temperature aging. The reason the room temperature evolution it is not obvious from the plots is due to the fact that common scales were chosen for the vertical (strain) axes in all of graphs in Figures 11-14.

10-4 10-5 10-6 10-7 10-8 Aging at Aging at Aging at Aging at

10-9 10-10 0

1

2

3

RT 75 oC o 100 C o 125 C

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Aging Time (month)

Figure 17 - Evolution of Creep Strain Rate with Aging (SAC305; Aging at 25, 75, 100 125 oC)


10-1

For the same stress level, Figures 15-18 and Figure 20 illustrate that the SAC alloys begin with creep rates lower than Sn-Pb solder immediately after reflow (zero aging). However, the changes in creep rate are much larger for the SAC alloys with elevated temperature aging relative to those experienced by Sn-Pb under the same conditions. Thus, “cross-over points” will occur where the SAC alloys begin to creep faster that Sn-Pb after certain durations of aging. These cross over points are illustrated in Figures 21-24. In these plots, the shaded regions represent the total extent of the creep rate variations for the Sn-Pb and SAC alloys. The crossovers occur most quickly for SAC105 and SAC205, but eventually also occur for SAC305 and SAC405 with longer durations of aging. For the stress level under consideration, crossovers did not occur for room temperature aging, which is similar to our previous results [49]. In general, the amount of aging required to obtain the creep rate cross-over depends in a complicated manner on the SAC alloy under consideration, the reflow profiles utilized, the applied stress level, and the aging temperature.

SAC405, RF o

T = 25 C

10-2

σ =15 MPa

Strain Rate (sec-1)


10-9


10-10 0

1

2

3

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5

Aging Time (month)

Figure 18 - Evolution of Creep Strain Rate with Aging (SAC405; Aging at 25, 75, 100 125 oC)

10-1

σ =15 MPa

10-4

Tin-Lead Creep Range

10-5 10-6 10-7 10-8 Aging Aging Aging Aging

10-9

0

1

2

3

4

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Aging Time (month)

Figure 21 - Creep Strain Rate Crossover Points for SAC105 and Sn-Pb 10 -1


σ =15 MPa

Cross-Over Points

63Sn-37Pb, RF

SAC205 Creep Range

10 -3

T = 25 oC

10-2

at RT o at 75 C o at 100 C at 125 oC

10-10

10 -2

10-1

SAC105 Creep Range

10-3

Comparison of SAC Aging to Sn-Pb Aging As observed in our previous studies [48-50], the aging induced changes in the creep strain rates of the SAC alloys are much larger than the analogous changes observed for conventional eutectic 63Sn-37Pb solder. Figure 20 contains the measured creep rate evolution curves for the Sn-Pb solder samples at the various aging temperatures. It is observed that the creep rates for Sn-Pb are restricted to a very narrow range of values between ε& = 1 x 10-6 and ε& = 1 x 10-5.

-1

Strain Rate (sec )

σ =15 MPa

10-3 -1

T = 25 oC,

Cross-Over Points

Figure 19 - Ratio of Maximum Creep Rate to Non-Aged Creep Rate for Each SAC Alloy

Strain Rate (sec )

SAC105, RF

10-2

-1

SAC105 SAC205 SAC305 SAC405

Maximum Aging Induced Increase in Creep Rate 5500X 1000X 220X 220X

Strain Rate (sec )

SAC Alloy

10-4 10-5

10 -4


10 -5 10 -6 10 -7

10-6

10 -8

10-7

10 -9

Aging Aging Aging Aging

at RT at 75 oC at 100 oC o at 125 C

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10-8

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Aging Time (month)

Figure 22 - Creep Strain Rate Crossover Points for SAC205 and Sn-Pb

5

Aging Time (month)

Figure 20 - Evolution of Creep Strain Rate with Aging (63Sn-37Pb; Aging at 25, 75, 100 125 oC)

108


10-1

position of Sn-Pb changes, the relative positions of the various SAC alloys remain fixed in eq. (2).


10-2

σ = 15 MPa

Cross-Over Points

SAC305 Creep Range

10-4

10-1


10-5 10-6

10-3 -1

10-7 10-8 Aging Aging Aging Aging

10-9 10-10 0

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RT o 75 C o 100 C o 125 C

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Aging Temperature: 25 oC σ = 15 MPa

10-2

Strain Rate (sec )

-1

Strain Rate (sec )

10-3

5

10-4 10-5

Increasing Silver Content

10-6 10-7

Aging Time (month) 10-8



10-9 10-10 0

T = 25 oC

σ = 15 MPa

Cross-Over Points

10-3

3

4

5

Figure 25 - SAC Creep Rate Comparisons (Aging at 25 oC)

SAC405 Creep Range

10-4

10-1 Aging Temperature: 75 oC


10-5 10-6

10-2

σ = 15 MPa

10-3 10-7

Strain Rate (sec-1)

-1

2

SAC405, RF

10-2

Strain Rate (sec )

1

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10-1

10-8 Aging at RT Aging at 75 oC Aging at 100 oC o Aging at 125 C

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10-4


10-5 10-6 10-7 10-8



10-9 10-10 0

ε& SAC105 > ε& SAC 205 > ε& SAC305 > ε& SAC 405

(2)

This expression is illustrated graphically in Figures 25-28, where the data from Figures 15-18 has been re-organized so that each plot is for a fixed aging temperature, and includes the results for all four of the SAC alloys. As mentioned above, it is not possible to insert the creep rate of Sn-Pb into the expression in eq. (2), since crossover points exist. With no aging, Sn-Pb has the highest creep rate and is at the left of the expression in eq. (2). However, after significant aging durations at elevated temperatures, Sn-Pb has the lowest creep rate and is at the right of the expression in eq. (2). While the

109

1

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Aging Time (month)

Figure 26 - SAC Creep Rate Comparisons (Aging at 75 oC) 10-1

o

Aging Temperature: 100 C σ = 15 MPa

10-2 10-3

Stain Rate (sec-1)

Comparisons of the Alloys at Each Aging Temperature In prior investigations, it has been observed that lowering of the silver content of a SAC alloy leads to increases in the creep rate [27, 31]. However, these prior studies did not examine the effects of silver content on aging. The data presented in this paper also suggest that lowering of the silver content of a SAC alloy leads to increases in the creep rate for all aging conditions. That is, the creep rates for a given applied stress, aging temperature, and aging time are ordered by:


10-4 10-5 10-6 10-7 10-8


10-9 10-10 0

1

2

3

4

5

Aging Time (month)

Figure 27 - SAC Creep Rate Comparisons (Aging at 100 oC)


10-1

Aging Temperature: 125 oC σ = 15 MPa

10-2

Strain Rate (sec-1)

10-3


10-4 10-5 10-6 10-7 10-8


10-9 10-10 0

1

2

3

4

5

Aging Time (month)

Figure 28 - SAC Creep Rate Comparisons (Aging at 125 oC) Effects of Elevated Temperature Aging on Microstructure Microstructure evolution during aging is the underlying reason for the changes in the creep properties of the solder alloys. For example, Figure 29 illustrates microstructural changes that occur in SAC105 solder during aging at 125 oC. The typical microstructure of SAC alloys consists of a Sn matrix (dendrites), and Ag3Sn and Cu6Sn5 second phases (intermetallics). During aging, the dendrites grow larger and merge, while at the same time the second phases develop into much larger particles (often needle-like in shape). This coarsening of the second phase particles is caused by diffusion, and is exacerbated at higher aging temperatures and longer aging durations. The diffusion rate of Ag and Cu at elevated temperatures will be much higher than at room temperature. Coarsened second phases will not be able to as effectively block dislocation movements, and thus there is a resulting loss of strength. At the same time the large secondary phases themselves become weak points in the materials. The creep deformation mechanisms for solders are mainly dislocation creep and grain sliding. The coarsened particles will also lose the ability to block grain boundary sliding. Thus the coarsening of the secondary phase particles leads to a dramatic loss of creep resistance in SAC alloys during aging. We have several parallel studies underway examining microstructural evolution during thermal aging of both unloaded solder specimens and of solder joints in various area array components (which are also subjected to stress during aging).

Summary and Conclusions The effects of aging on mechanical behavior of lead free solders have been examined by performing creep tests on four different SAC alloys (SAC105, SAC205, SAC305, SAC405) that were aged for various durations (0-4 months) at room temperature (25 oC), and several elevated temperatures (75, 100, and 125 oC). Analogous tests were performed with 63Sn-37Pb eutectic solder samples for comparison purposes. Variations of the creep properties were observed and modeled as a function of aging time and aging temperature. In addition, the chosen selection of SAC alloys has allowed us to explore the effects of silver content on aging behavior (we have examined SACN05 with N= 1%, 2%, 3%, and 4% silver; with all alloys containing 0.5% copper). The results obtained in this work have demonstrated the significant effects of elevated temperature exposure on the creep behavior of solder joints. As expected, the creep rates evolved more dramatically when the aging temperature was increased. In addition, the effects of aging were shown to be significant even for aging temperature slightly above room temperature (e.g. T = 75 oC). The recorded data demonstrate that the creep behaviors of lead free and tin-lead solders experience a “cross-over point” where the lead free solders begin to creep at higher rates than standard 63Sn-37Pb solder for the same stress level. These cross-over points were observed to exist for all four of the SAC solder alloys, and for all of the aging temperatures except for room temperature. In addition, the lower silver content alloys (e.g. SAC105) were observed to be more much more sensitive to aging (have greater changes in the creep rate for a given aging time) than the higher silver content alloys (e.g. SAC405). The times required before the cross-over occurred were reduced when considering higher aging temperatures or SAC alloys with lower silver content. It was also observed that lowering of the silver content of a SAC alloy leads to increases in the creep rates for all aging conditions. The degradations of the mechanical and creep properties of lead free SAC solders during aging are caused by microstructural evolution. In particular, there is dramatic coarsening of the secondary intermetallic particles. When the particles are small and fine precipitations, they can effectively block the movement of dislocations and reduce grain boundaries sliding, thus strengthening the materials and enhancing creep resistance. When the second phase particles grow larger, their ability to block the dislocation movements and grain boundary sliding are significantly reduced leading to reduced strength and to degraded resistance to creep deformations. Acknowledgments This work was supported by Basic Research Grant # IIP0434909-015 from the National Science Foundation, as well as the NSF Center for Advanced Vehicle Electronics (CAVE).

Figure 29 - Microstructural Evolution During of SAC105 Solder Subjected to Aging at 125 oC

110


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