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Failure mode analysis of lead-free solder joints under high speed impact testing
Materials Science and Engineering A 494 (2008) 196–202
Contents lists available at ScienceDirect
Materials Science and Engineering A journal homepage: www.elsevier.com/locate/msea
Failure mode analysis of lead-free solder joints under high speed impact testing De-Shin Liu a , Chia-Yuan Kuo a,∗ , Chang-Lin Hsu a , Geng-Shin Shen b , Yu-Ren Chen b , Kuo-Cheng Lo b a b
National Chung Cheng University, Department of Mechanical Engineering, Chia-Yi, Taiwan, ROC ChipMOS TECHNOLOGIES LTD, Tainan, Taiwan, ROC
a r t i c l e
i n f o
Article history: Received 8 November 2007 Received in revised form 7 April 2008 Accepted 8 April 2008 Keywords: Lead-free solder Micro-impact test Failure mode analysis
a b s t r a c t Using an Instron micro-impact system, this study investigates the failure characteristics of 96.5Sn–3Ag–0.5Cu lead-free solder joints aged at either room temperature or 125 ◦ C, respectively, and then impacted at shear rates of up to 1 m/s. Four types of failure mode are identified, namely M1: interfacial fracture with no residual solder left on the pad; M2: interfacial fracture with residual solder left on the pad; M3: solder ball fracture; and M4: substrate fracture. The experimental results reveal that the solder specimens fail in different failure modes at the same impact speed. The transition from ductile to brittle failure occurs at an impact speed of around 0.5 m/s. At an impact speed of 0.7 (±0.05) m/s or more, over 70% of the specimens fail in the M1 or M2 modes under all of the testing conditions. The isothermal aging process is found to reduce the interfacial strength, and hence the percentage of M3 and M4 mode failures reduces significantly. Overall, the experimental results suggest that the failure mode distribution obtained in high speed impact tests performed at 0.5 m/s provides a feasible component-level quality assurance index. © 2008 Elsevier B.V. All rights reserved.
1. Introduction In response to health and safety concerns, the packaging of electronic products is now performed exclusively using lead-free solder joints. Although various techniques have been proposed for assessing the mechanical strength of the solder ball bonds in ball grid array (BGA) packages, the ball shear test defined in standard JESD22-B117 is one of the most common [1]. Huang et al. constructed a two-dimensional finite element model of the solder ball shear test configuration and investigated the effect of the shear height and the shear speed on the shear force–displacement curve. The results indicated that the shear force–displacement curve trend increased with a faster shear speed and a lower shear height. The numerical results were found to be consistent with the experimental data and led the authors to conclude that ball shear tests should be conducted at a shear height of less than 25% of the solder ball height and a shear speed slower than 200 m/s [2]. Canumalla et al. performed a series of experimental tests to examine the packageto-board interconnection shear strength of solder balls with various surface finishes [3]. Chia et al. reported that in ball shear tests, the shear strength, total shear work and shear work up to maximum load all increase with an increasing shear rate [4].
∗ Corresponding author. Fax: +886 5 2720589. E-mail address: [email protected] (C.-Y. Kuo). 0921-5093/$ – see front matter © 2008 Elsevier B.V. All rights reserved. doi:10.1016/j.msea.2008.04.015
Portable electronic products most commonly fail as a result of being accidentally dropped to the floor, resulting in a breakage of the solder interconnect. Therefore, developing reliable methods for measuring the impact loading capacity of electronic packages is an essential concern. Data et al. performed a series of Charpy impact tests to evaluate the effect of aging on the impact reliability of SnZn solder joints. The results revealed that the joints bonded on electroless Au/Ni(P) pads exhibited a ductile failure mode and independent of the aging time [5]. Using the same technique, Ou et al. showed that the impact toughness of solder joints improved following an isothermal aging process when the types of solder Sn1Ag, Sn1Ag0.5Cu and Sn1Ag0.5Cu1In bonded on Au/electrolytic Ni/Cu pad [6]. Wong et al. investigated the effects of the pad finish, the solder mask design and the isothermal aging conditions on the mechanical behavior of various solders. The results demonstrated that the use of a solder mask yielded a significant improvement in the impact strength of the solder ball. In general, the impact strength of solder joints and their ability to absorb impact energy are crucial factors in determining the survivability of the boardlevel interconnections of a portable electronic device in the event of a sudden impact. As a result, high speed shear tests play a key role in predicting the component-level quality of electronic products [7]. Newman found that both the shear strength and the pull strength of solder joints increased with an increasing test speed. Consequently, the solder joints tended to fracture at the pad interface when subjected to high speed testing [8]. Yeh and Lai presented a numerical approach for correlating the results of a ball impact test
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Fig. 3. Solder reflow curve.
Fig. 1. Photograph of experimental setup.
Fig. 4. Photograph showing test sample adhered to steel plate.
Fig. 2. Schematic illustration of the solder joint specimen.
with the board-level drop reliability of portable electronic components [9]. The majority of the studies outlined above focus primarily on the effect of the impact speed on the mechanical response of the solder joint. However, reviewing the literature, it appears that very few studies have attempted a systematic investigation into the type and frequency of the different failure modes which may occur at any given test speed. Accordingly, the current study conducts an extensive experimental investigation into the failure characteristics of lead-free 96.5Sn–3Ag–0.5Cu solder balls under high speed impacts. Using a high speed impact test system, force–displacement curves are generated for specimens aged at room temperature or at 125 ◦ C for up to 720 h and then impacted at velocities ranging from 0.3 to 1.0 m/s. The fractured specimens are observed using optical microscopy (OM) and scanning electron microscopy (SEM) techniques. The experimental results are then analyzed in order to evaluate the respective effects of the impact speed and the aging
Fig. 5. Illustration of four failure modes.
conditions on the failure mode, the maximum peak load and the energy to peak load of the various specimens. 2. Experimental procedure The experimental tests were performed using an Instron microimpact system capable of generating impact speeds as high as 1.0 m/s. As shown in Fig. 1, the main hardware components of the experimental setup included a striking head assembly, a veloc-
Table 1 Impact velocity and aging parameters Impact speed (m/s) Room temperature Baking (125 ◦ C) Aging time index No. 1 Hours After reflow
0.3
2 12
3 24
0.5
4 48
5 72
6 96
0.7
7 120
8 168
1.0 1.0
9 216
10 384
11 552
12 720
198
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Fig. 6. (a) OM and (b) SEM images of four failure modes.
ity control module, a specimen fixture, and an XYZ stage. The load was sensed using a 45 N load cell with a linearity of 0.4% FS and was acquired continuously over the course of the impact. The displacement was measured using a linear variable differen-
tial transformer (LVDT) with a measurement resolution of 0.1 m. A vision alignment system comprising two charge coupled device (CCD) cameras fitted with microscope lenses were used to check the parallel alignment of the striking head with the specimen
Fig. 7. Failure mode distribution of RT specimens under impact velocities of: (a) 0.3 m/s, (b) 0.5 m/s, (c) 0.7 m/s and (d) 1.0 m/s.
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speeds of 0.3, 0.5, 0.7 and 1.0 (±0.05) m/s, respectively, while the isothermally aged specimens were tested only at the highest speed of 1.0 (±0.05) m/s. Both sets of specimens were aged for intervals of between 0 and 720 h prior to testing. As shown in the lower row of Table 1, a total of 12 different aging intervals were considered. To ensure the reliability of the experimental results, 66 solder balls were tested under each experimental condition. Following the impact tests, the specimens were observed using OM and SEM techniques. 3. Experimental results and discussion
Fig. 8. Failure mode distribution of baked specimens under impact velocity of 1.0 m/s.
prior to testing. As shown in Fig. 1, the entire test apparatus was placed on an optical table to minimize the effects of environmental noise. The test specimen comprised a bismaleimide triazine (BT) laminate substrate measuring 12.5 mm × 5 mm coated with an electrolytic nickel/gold (Ni/Au) pad finish. A row of 11 solder balls of diameter 450 m was deposited on the substrate using a solder mask with an opening of 380 m, the schematic illustration of the solder joint specimen was shown in Fig. 2. Fig. 3 illustrates the reflow profile of the solder balls. Following the reflow process, half of the specimens were placed in an oven and aged isothermally at a temperature of 125 ◦ C, while the remainders were left to age naturally at room temperature. Prior to impact testing, the specimens were adhered to a steel plate measuring 18 mm × 10 mm (see Fig. 4) and were then mounted in the specimen fixture. The impact height was adjusted to 50 m using the XYZ stage and verified using the vision alignment system. A series of experiments was performed to evaluate the effect of the impact speed and the aging time on the failure mode of the various specimens. As shown in Table 1, the specimens aged at room temperature (designated hereafter as the RT specimens) were tested at impact
Fig. 9. Variation of impact load over impact time.
The OM and SEM observations revealed that the specimens failed in one of four different failure modes, as illustrated schematically in Fig. 5. In failure mode 1 (designated hereafter as M1), the fracture occurred at the interface between the ball and the pad and no residual solder was left on the pad. In the second failure mode (M2), the specimen fractured at the interface between the ball and the pad and some residual solder remained on the pad. Both modes M1 and M2 are classified as brittle failure. In the third failure mode (M3), the fracture occurred within the solder ball itself. Finally, in the fourth failure mode (M4), the fracture occurred in the substrate, resulting in the removal of the pad. Both modes M3 and M4 are classified as ductile failure. Fig. 6(a) and (b) present OM and SEM images of the four failure modes. Figs. 7 and 8 present the failure mode distributions of the RT and isothermally aged specimens, respectively. The results presented in Fig. 7 indicate that for impact speeds of 0.7 (±0.05) m/s or more, failure modes M1 and M2 account for
Fig. 10. Test results for M1 specimens: (a) peak load versus time and (b) energy to peak load versus time.
200
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more than 70% of the total specimen failures. From inspection, it can be seen that the frequency of the M3 and M4 failures reduces as the impact speed increases, while that of the M1 and M2 failures increases. This trend indicates that the joint failure changes from a ductile mode to a brittle mode as the impact speed is increased. Comparing Figs. 7(d) and 8, it can be seen that the annealing process yields a significant reduction in the occurrence of M3 and M4 failures. Fig. 9 illustrates the variation of the impact load over the impact time for specimens tested at a speed of 1 m/s. According to Chia et al. [4], the shear work performed after the point of maximum load is insensitive to the shear rate and the solder toughness. As a result, the energy absorption capacity of the solder specimens is defined as the area under the force–displacement curve before the point of peak load. Fig. 10(a) and (b) illustrates the effect of the aging time and the impact speed on the peak load and energy to peak load characteristics of the RT and baked (BK) specimens which fail with an M1 failure mode. The results presented in Fig. 10(a) show that at impact speeds lower than 0.7 m/s, the peak load is decreases slightly as the aging time more than 168 h (i.e. 7 days). In the impact tests performed at the highest speed of 1 m/s, the peak load is insensitive to the aging time. Comparing the RT and BK profiles, it can be seen that the annealing process results in a significant reduction in the peak load. Fig. 10(b) shows that the energy to peak load of the fractured specimens is relatively insensitive to both the impact speed and the aging time.
Fig. 12. Test results for M3 specimens: (a) peak load versus time and (b) energy to peak load versus time.
Fig. 11. Test results for M2 specimens: (a) peak load versus time and (b) energy to peak load versus time.
Fig. 11(a) and (b) shows the peak load and energy to peak load characteristics of the RT and BK specimens which experience an M2 failure. The results show that at impact speeds greater than 0.5 m/s, the peak load and energy to peak load both remain approximately constant when the aging time increases beyond 168 h. It can also be seen that the specimens impacted at the lowest speed of 0.3 m/s have a lower peak load and a greater energy absorption capacity than those deformed at higher speeds. Moreover, the level of the peak load at various impact rates could be differentiated clearly after168 h of aging. The results again show that the isothermal aging process leads to a notable reduction in the peak load of the impacted specimens. Fig. 12(a) and (b) shows that for the RT and BK specimens which experience an M3 failure, the peak load and the energy to peak load reduce significantly when the aging time is increased to 168 h, but then remain relatively constant when the aging time is increased further at impact speeds 0.3 and 0.5 m/s; there is no noticeable trend at the impact speeds 0.7 and 1.0 m/s. Furthermore, it can be seen that the peak load increases with an increasing impact speed, while the energy to peak load decreases. As a result, the energy absorption capacity of the specimen impacted at 0.3 m/s is almost twice that of the specimen deformed at 1.0 m/s. In contrast to the M1 and M2 specimens, it can be seen that the isothermal aging process yields a less significant reduction in the peak load of the M3 specimens. Fig. 13(a) and (b) presents the peak load and energy to peak load characteristics of the M4 specimens. Since relatively few of the current specimens actually failed with this particular failure
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Fig. 13. Test results for M4 specimens: (a) peak load versus time and (b) energy to peak load versus time.
mode (see Figs. 7 and 8), it is difficult to draw any reliable conclusions regarding the effects of the impact speed and the aging time. However, it is noted that both the peak load and the energy to peak load are generally lower than the corresponding values for the M3 specimens. Table 2 summarizes the overall effects of the impact speed and the aging time on the peak load and energy to peak load of the various RT and BK specimens. Meanwhile, Table 3 presents the peak load and energy to peak load values of the RT specimens at the aging time reach 168 h (i.e. 7 days). From inspection, it is found that the four failure modes can be ranked in terms of a decreasing energy to peak load as follows: M3 (solder ball fracture) > mode M4 (substrate
Table 2 Summary of impact speed and aging condition effects on peak load and energy to peak load characteristics of failed specimens
Table 3 Peak load and energy to peak load characteristics of failed specimens aged for 168 h at room temperature Failure mode M1
M2
M3
M4
V = 0.3 m/s Peak load (N) Energy to peak load (mJ)
9.5 0.4
9.6 0.6
11.1 1.7
11.0 0.9
V = 0.5 Peak load (N) Energy to peak load (mJ)
10.6 0.3
9.5 0.3
11.7 1.2
11.1 0.6
V = 0.7 Peak load (N) Energy to peak load (mJ)
9.8 0.2
9.4 0.3
12.0 1.1
12.7 0.7
V = 1.0 Peak load (N) Energy to peak load (mJ)
8.5 0.2
9.4 0.3
12.9 0.9
12.9 0.6
fracture) > mode M2 (interfacial failure with residual solder on the pad) > mode M1 (interfacial failure with no residual solder on the pad). 4. Conclusions This study has performed an experimental investigation into the effects of the impact speed and aging time on the strength and failure modes of 96.5Sn–3Ag–0.5Cu lead-free solder joints under high speed impacts ranging from 0.3 to 1.0 m/s. The experimental results support the following conclusions: • High speed shear impacts induce different failure modes under the same impact speed. Consequently, a large number of test samples are required to establish reliable estimates of the failure mode distribution. Moreover, failure mode percentage would be an equal important quality assurance index as peak load and energy to peak load. • Four failure modes may occur at any given impact speed, M1: interfacial fracture with no residual solder on the pad; M2: interfacial fracture with residual solder on the pad; M3: solder ball fracture; and M4: substrate fracture. • The transition from a ductile failure mode to a brittle failure mode occurs at an impact speed of around 0.5 m/s. For an impact speed
202
•
•
•
• •
•
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of 0.7 (±0.05) m/s or more, failure modes M1 and M2 account for over 70% of all the failures in all the testing conditions. The isothermal aging process conducted at a temperature of 125 ◦ C yields a significant reduction in the occurrence of M3 and M4 failures. For a given impact velocity, the failure modes can be ranked in terms of a reducing energy absorption capacity as follows: M3 > M4 > M2 > M1. For the M3 specimens, the solder strength increases with an increasing impact speed, whereas the energy absorption capacity decreases. However, in the M2 specimens, the interfacial strength and peak energy are insensitive to the impact speed at speeds greater than 0.5 m/s. The isothermal aging process results in a reduction in the interfacial strength of the solder joint. 96.5Sn–3Ag–0.5Cu lead-free solder joints should be aged for 168 h following the reflow process in order to get stable trend of impact strength and energy. The failure mode distribution obtained from high speed impact tests performed at an impact rate of 0.5 m/s provides a feasible component-level quality assurance index.
Acknowledgement The current authors gratefully acknowledge the financial support provided to this study by the National Science Council, Taiwan, ROC, under Grant No. NSC96-2221-E-194-026. References [1] JEDEC, JESD22-B117 (2000), http://www.jedec.org. [2] X. Huang, S.-W.R. Lee, C.C. Yuan, S. Hui, Proceedings of the 51th Electronic Components and Technology Conference, IEEE, Piscataway, NJ, 2001, pp. 1065–1071. [3] S. Canumalla, H.-D. Yang, P. Viswanadham, T.O. Reinikainen, IEEE Trans. Comp. Packag. Technol. 27 (2004) 182–190. [4] J.Y.H. Chia, B. Cotterell, T.C. Chai, Mater. Sci. Eng. A 417 (2006) 259–274. [5] M. Date, T. Shoji, M. Fujiyoshi, K. Sato, K.N. Tu, Scr. Mater. 51 (2004) 641–645. [6] S. Ou, Y. Xu, K.N. Tu, M.O. Alam, Y.C. Chan, Proceedings of the 55th Electronic Components and Technology Conference, IEEE, NJ, 2005, pp. 467–471. [7] E.H. Wong, R. Rajoo, Y.W. Mai, S.K.W. Seah, K.T. Tsai, L.M. Yap, Proceedings of the 55th Electronic Components and Technology Conference, IEEE, NJ, 2005, pp. 1202–1209. [8] K. Newman, Proceedings of the 55th Electronic Components and Technology Conference, IEEE, NJ, 2005, pp. 1194–1201. [9] C.-L. Yeh, Y.-S. Lai, J. Electron. Mater. 35 (2006) 1892–1901.
Contents lists available at ScienceDirect
Materials Science and Engineering A journal homepage: www.elsevier.com/locate/msea
Failure mode analysis of lead-free solder joints under high speed impact testing De-Shin Liu a , Chia-Yuan Kuo a,∗ , Chang-Lin Hsu a , Geng-Shin Shen b , Yu-Ren Chen b , Kuo-Cheng Lo b a b
National Chung Cheng University, Department of Mechanical Engineering, Chia-Yi, Taiwan, ROC ChipMOS TECHNOLOGIES LTD, Tainan, Taiwan, ROC
a r t i c l e
i n f o
Article history: Received 8 November 2007 Received in revised form 7 April 2008 Accepted 8 April 2008 Keywords: Lead-free solder Micro-impact test Failure mode analysis
a b s t r a c t Using an Instron micro-impact system, this study investigates the failure characteristics of 96.5Sn–3Ag–0.5Cu lead-free solder joints aged at either room temperature or 125 ◦ C, respectively, and then impacted at shear rates of up to 1 m/s. Four types of failure mode are identified, namely M1: interfacial fracture with no residual solder left on the pad; M2: interfacial fracture with residual solder left on the pad; M3: solder ball fracture; and M4: substrate fracture. The experimental results reveal that the solder specimens fail in different failure modes at the same impact speed. The transition from ductile to brittle failure occurs at an impact speed of around 0.5 m/s. At an impact speed of 0.7 (±0.05) m/s or more, over 70% of the specimens fail in the M1 or M2 modes under all of the testing conditions. The isothermal aging process is found to reduce the interfacial strength, and hence the percentage of M3 and M4 mode failures reduces significantly. Overall, the experimental results suggest that the failure mode distribution obtained in high speed impact tests performed at 0.5 m/s provides a feasible component-level quality assurance index. © 2008 Elsevier B.V. All rights reserved.
1. Introduction In response to health and safety concerns, the packaging of electronic products is now performed exclusively using lead-free solder joints. Although various techniques have been proposed for assessing the mechanical strength of the solder ball bonds in ball grid array (BGA) packages, the ball shear test defined in standard JESD22-B117 is one of the most common [1]. Huang et al. constructed a two-dimensional finite element model of the solder ball shear test configuration and investigated the effect of the shear height and the shear speed on the shear force–displacement curve. The results indicated that the shear force–displacement curve trend increased with a faster shear speed and a lower shear height. The numerical results were found to be consistent with the experimental data and led the authors to conclude that ball shear tests should be conducted at a shear height of less than 25% of the solder ball height and a shear speed slower than 200 m/s [2]. Canumalla et al. performed a series of experimental tests to examine the packageto-board interconnection shear strength of solder balls with various surface finishes [3]. Chia et al. reported that in ball shear tests, the shear strength, total shear work and shear work up to maximum load all increase with an increasing shear rate [4].
∗ Corresponding author. Fax: +886 5 2720589. E-mail address: [email protected] (C.-Y. Kuo). 0921-5093/$ – see front matter © 2008 Elsevier B.V. All rights reserved. doi:10.1016/j.msea.2008.04.015
Portable electronic products most commonly fail as a result of being accidentally dropped to the floor, resulting in a breakage of the solder interconnect. Therefore, developing reliable methods for measuring the impact loading capacity of electronic packages is an essential concern. Data et al. performed a series of Charpy impact tests to evaluate the effect of aging on the impact reliability of SnZn solder joints. The results revealed that the joints bonded on electroless Au/Ni(P) pads exhibited a ductile failure mode and independent of the aging time [5]. Using the same technique, Ou et al. showed that the impact toughness of solder joints improved following an isothermal aging process when the types of solder Sn1Ag, Sn1Ag0.5Cu and Sn1Ag0.5Cu1In bonded on Au/electrolytic Ni/Cu pad [6]. Wong et al. investigated the effects of the pad finish, the solder mask design and the isothermal aging conditions on the mechanical behavior of various solders. The results demonstrated that the use of a solder mask yielded a significant improvement in the impact strength of the solder ball. In general, the impact strength of solder joints and their ability to absorb impact energy are crucial factors in determining the survivability of the boardlevel interconnections of a portable electronic device in the event of a sudden impact. As a result, high speed shear tests play a key role in predicting the component-level quality of electronic products [7]. Newman found that both the shear strength and the pull strength of solder joints increased with an increasing test speed. Consequently, the solder joints tended to fracture at the pad interface when subjected to high speed testing [8]. Yeh and Lai presented a numerical approach for correlating the results of a ball impact test
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Fig. 3. Solder reflow curve.
Fig. 1. Photograph of experimental setup.
Fig. 4. Photograph showing test sample adhered to steel plate.
Fig. 2. Schematic illustration of the solder joint specimen.
with the board-level drop reliability of portable electronic components [9]. The majority of the studies outlined above focus primarily on the effect of the impact speed on the mechanical response of the solder joint. However, reviewing the literature, it appears that very few studies have attempted a systematic investigation into the type and frequency of the different failure modes which may occur at any given test speed. Accordingly, the current study conducts an extensive experimental investigation into the failure characteristics of lead-free 96.5Sn–3Ag–0.5Cu solder balls under high speed impacts. Using a high speed impact test system, force–displacement curves are generated for specimens aged at room temperature or at 125 ◦ C for up to 720 h and then impacted at velocities ranging from 0.3 to 1.0 m/s. The fractured specimens are observed using optical microscopy (OM) and scanning electron microscopy (SEM) techniques. The experimental results are then analyzed in order to evaluate the respective effects of the impact speed and the aging
Fig. 5. Illustration of four failure modes.
conditions on the failure mode, the maximum peak load and the energy to peak load of the various specimens. 2. Experimental procedure The experimental tests were performed using an Instron microimpact system capable of generating impact speeds as high as 1.0 m/s. As shown in Fig. 1, the main hardware components of the experimental setup included a striking head assembly, a veloc-
Table 1 Impact velocity and aging parameters Impact speed (m/s) Room temperature Baking (125 ◦ C) Aging time index No. 1 Hours After reflow
0.3
2 12
3 24
0.5
4 48
5 72
6 96
0.7
7 120
8 168
1.0 1.0
9 216
10 384
11 552
12 720
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Fig. 6. (a) OM and (b) SEM images of four failure modes.
ity control module, a specimen fixture, and an XYZ stage. The load was sensed using a 45 N load cell with a linearity of 0.4% FS and was acquired continuously over the course of the impact. The displacement was measured using a linear variable differen-
tial transformer (LVDT) with a measurement resolution of 0.1 m. A vision alignment system comprising two charge coupled device (CCD) cameras fitted with microscope lenses were used to check the parallel alignment of the striking head with the specimen
Fig. 7. Failure mode distribution of RT specimens under impact velocities of: (a) 0.3 m/s, (b) 0.5 m/s, (c) 0.7 m/s and (d) 1.0 m/s.
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speeds of 0.3, 0.5, 0.7 and 1.0 (±0.05) m/s, respectively, while the isothermally aged specimens were tested only at the highest speed of 1.0 (±0.05) m/s. Both sets of specimens were aged for intervals of between 0 and 720 h prior to testing. As shown in the lower row of Table 1, a total of 12 different aging intervals were considered. To ensure the reliability of the experimental results, 66 solder balls were tested under each experimental condition. Following the impact tests, the specimens were observed using OM and SEM techniques. 3. Experimental results and discussion
Fig. 8. Failure mode distribution of baked specimens under impact velocity of 1.0 m/s.
prior to testing. As shown in Fig. 1, the entire test apparatus was placed on an optical table to minimize the effects of environmental noise. The test specimen comprised a bismaleimide triazine (BT) laminate substrate measuring 12.5 mm × 5 mm coated with an electrolytic nickel/gold (Ni/Au) pad finish. A row of 11 solder balls of diameter 450 m was deposited on the substrate using a solder mask with an opening of 380 m, the schematic illustration of the solder joint specimen was shown in Fig. 2. Fig. 3 illustrates the reflow profile of the solder balls. Following the reflow process, half of the specimens were placed in an oven and aged isothermally at a temperature of 125 ◦ C, while the remainders were left to age naturally at room temperature. Prior to impact testing, the specimens were adhered to a steel plate measuring 18 mm × 10 mm (see Fig. 4) and were then mounted in the specimen fixture. The impact height was adjusted to 50 m using the XYZ stage and verified using the vision alignment system. A series of experiments was performed to evaluate the effect of the impact speed and the aging time on the failure mode of the various specimens. As shown in Table 1, the specimens aged at room temperature (designated hereafter as the RT specimens) were tested at impact
Fig. 9. Variation of impact load over impact time.
The OM and SEM observations revealed that the specimens failed in one of four different failure modes, as illustrated schematically in Fig. 5. In failure mode 1 (designated hereafter as M1), the fracture occurred at the interface between the ball and the pad and no residual solder was left on the pad. In the second failure mode (M2), the specimen fractured at the interface between the ball and the pad and some residual solder remained on the pad. Both modes M1 and M2 are classified as brittle failure. In the third failure mode (M3), the fracture occurred within the solder ball itself. Finally, in the fourth failure mode (M4), the fracture occurred in the substrate, resulting in the removal of the pad. Both modes M3 and M4 are classified as ductile failure. Fig. 6(a) and (b) present OM and SEM images of the four failure modes. Figs. 7 and 8 present the failure mode distributions of the RT and isothermally aged specimens, respectively. The results presented in Fig. 7 indicate that for impact speeds of 0.7 (±0.05) m/s or more, failure modes M1 and M2 account for
Fig. 10. Test results for M1 specimens: (a) peak load versus time and (b) energy to peak load versus time.
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more than 70% of the total specimen failures. From inspection, it can be seen that the frequency of the M3 and M4 failures reduces as the impact speed increases, while that of the M1 and M2 failures increases. This trend indicates that the joint failure changes from a ductile mode to a brittle mode as the impact speed is increased. Comparing Figs. 7(d) and 8, it can be seen that the annealing process yields a significant reduction in the occurrence of M3 and M4 failures. Fig. 9 illustrates the variation of the impact load over the impact time for specimens tested at a speed of 1 m/s. According to Chia et al. [4], the shear work performed after the point of maximum load is insensitive to the shear rate and the solder toughness. As a result, the energy absorption capacity of the solder specimens is defined as the area under the force–displacement curve before the point of peak load. Fig. 10(a) and (b) illustrates the effect of the aging time and the impact speed on the peak load and energy to peak load characteristics of the RT and baked (BK) specimens which fail with an M1 failure mode. The results presented in Fig. 10(a) show that at impact speeds lower than 0.7 m/s, the peak load is decreases slightly as the aging time more than 168 h (i.e. 7 days). In the impact tests performed at the highest speed of 1 m/s, the peak load is insensitive to the aging time. Comparing the RT and BK profiles, it can be seen that the annealing process results in a significant reduction in the peak load. Fig. 10(b) shows that the energy to peak load of the fractured specimens is relatively insensitive to both the impact speed and the aging time.
Fig. 12. Test results for M3 specimens: (a) peak load versus time and (b) energy to peak load versus time.
Fig. 11. Test results for M2 specimens: (a) peak load versus time and (b) energy to peak load versus time.
Fig. 11(a) and (b) shows the peak load and energy to peak load characteristics of the RT and BK specimens which experience an M2 failure. The results show that at impact speeds greater than 0.5 m/s, the peak load and energy to peak load both remain approximately constant when the aging time increases beyond 168 h. It can also be seen that the specimens impacted at the lowest speed of 0.3 m/s have a lower peak load and a greater energy absorption capacity than those deformed at higher speeds. Moreover, the level of the peak load at various impact rates could be differentiated clearly after168 h of aging. The results again show that the isothermal aging process leads to a notable reduction in the peak load of the impacted specimens. Fig. 12(a) and (b) shows that for the RT and BK specimens which experience an M3 failure, the peak load and the energy to peak load reduce significantly when the aging time is increased to 168 h, but then remain relatively constant when the aging time is increased further at impact speeds 0.3 and 0.5 m/s; there is no noticeable trend at the impact speeds 0.7 and 1.0 m/s. Furthermore, it can be seen that the peak load increases with an increasing impact speed, while the energy to peak load decreases. As a result, the energy absorption capacity of the specimen impacted at 0.3 m/s is almost twice that of the specimen deformed at 1.0 m/s. In contrast to the M1 and M2 specimens, it can be seen that the isothermal aging process yields a less significant reduction in the peak load of the M3 specimens. Fig. 13(a) and (b) presents the peak load and energy to peak load characteristics of the M4 specimens. Since relatively few of the current specimens actually failed with this particular failure
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Fig. 13. Test results for M4 specimens: (a) peak load versus time and (b) energy to peak load versus time.
mode (see Figs. 7 and 8), it is difficult to draw any reliable conclusions regarding the effects of the impact speed and the aging time. However, it is noted that both the peak load and the energy to peak load are generally lower than the corresponding values for the M3 specimens. Table 2 summarizes the overall effects of the impact speed and the aging time on the peak load and energy to peak load of the various RT and BK specimens. Meanwhile, Table 3 presents the peak load and energy to peak load values of the RT specimens at the aging time reach 168 h (i.e. 7 days). From inspection, it is found that the four failure modes can be ranked in terms of a decreasing energy to peak load as follows: M3 (solder ball fracture) > mode M4 (substrate
Table 2 Summary of impact speed and aging condition effects on peak load and energy to peak load characteristics of failed specimens
Table 3 Peak load and energy to peak load characteristics of failed specimens aged for 168 h at room temperature Failure mode M1
M2
M3
M4
V = 0.3 m/s Peak load (N) Energy to peak load (mJ)
9.5 0.4
9.6 0.6
11.1 1.7
11.0 0.9
V = 0.5 Peak load (N) Energy to peak load (mJ)
10.6 0.3
9.5 0.3
11.7 1.2
11.1 0.6
V = 0.7 Peak load (N) Energy to peak load (mJ)
9.8 0.2
9.4 0.3
12.0 1.1
12.7 0.7
V = 1.0 Peak load (N) Energy to peak load (mJ)
8.5 0.2
9.4 0.3
12.9 0.9
12.9 0.6
fracture) > mode M2 (interfacial failure with residual solder on the pad) > mode M1 (interfacial failure with no residual solder on the pad). 4. Conclusions This study has performed an experimental investigation into the effects of the impact speed and aging time on the strength and failure modes of 96.5Sn–3Ag–0.5Cu lead-free solder joints under high speed impacts ranging from 0.3 to 1.0 m/s. The experimental results support the following conclusions: • High speed shear impacts induce different failure modes under the same impact speed. Consequently, a large number of test samples are required to establish reliable estimates of the failure mode distribution. Moreover, failure mode percentage would be an equal important quality assurance index as peak load and energy to peak load. • Four failure modes may occur at any given impact speed, M1: interfacial fracture with no residual solder on the pad; M2: interfacial fracture with residual solder on the pad; M3: solder ball fracture; and M4: substrate fracture. • The transition from a ductile failure mode to a brittle failure mode occurs at an impact speed of around 0.5 m/s. For an impact speed
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of 0.7 (±0.05) m/s or more, failure modes M1 and M2 account for over 70% of all the failures in all the testing conditions. The isothermal aging process conducted at a temperature of 125 ◦ C yields a significant reduction in the occurrence of M3 and M4 failures. For a given impact velocity, the failure modes can be ranked in terms of a reducing energy absorption capacity as follows: M3 > M4 > M2 > M1. For the M3 specimens, the solder strength increases with an increasing impact speed, whereas the energy absorption capacity decreases. However, in the M2 specimens, the interfacial strength and peak energy are insensitive to the impact speed at speeds greater than 0.5 m/s. The isothermal aging process results in a reduction in the interfacial strength of the solder joint. 96.5Sn–3Ag–0.5Cu lead-free solder joints should be aged for 168 h following the reflow process in order to get stable trend of impact strength and energy. The failure mode distribution obtained from high speed impact tests performed at an impact rate of 0.5 m/s provides a feasible component-level quality assurance index.
Acknowledgement The current authors gratefully acknowledge the financial support provided to this study by the National Science Council, Taiwan, ROC, under Grant No. NSC96-2221-E-194-026. References [1] JEDEC, JESD22-B117 (2000), http://www.jedec.org. [2] X. Huang, S.-W.R. Lee, C.C. Yuan, S. Hui, Proceedings of the 51th Electronic Components and Technology Conference, IEEE, Piscataway, NJ, 2001, pp. 1065–1071. [3] S. Canumalla, H.-D. Yang, P. Viswanadham, T.O. Reinikainen, IEEE Trans. Comp. Packag. Technol. 27 (2004) 182–190. [4] J.Y.H. Chia, B. Cotterell, T.C. Chai, Mater. Sci. Eng. A 417 (2006) 259–274. [5] M. Date, T. Shoji, M. Fujiyoshi, K. Sato, K.N. Tu, Scr. Mater. 51 (2004) 641–645. [6] S. Ou, Y. Xu, K.N. Tu, M.O. Alam, Y.C. Chan, Proceedings of the 55th Electronic Components and Technology Conference, IEEE, NJ, 2005, pp. 467–471. [7] E.H. Wong, R. Rajoo, Y.W. Mai, S.K.W. Seah, K.T. Tsai, L.M. Yap, Proceedings of the 55th Electronic Components and Technology Conference, IEEE, NJ, 2005, pp. 1202–1209. [8] K. Newman, Proceedings of the 55th Electronic Components and Technology Conference, IEEE, NJ, 2005, pp. 1194–1201. [9] C.-L. Yeh, Y.-S. Lai, J. Electron. Mater. 35 (2006) 1892–1901.