Rate-dependent properties of Sn–Ag–Cu based lead-free solder joints for WLCSP
Y.A.Su a ,L.B.Tan a ,T.Y.Tee b ,V.B.C.Tan a,*
a National University of Singapore,Department of Mechanical Engineering,9Engineering Drive 1,Singapore 117576,Singapore b
Amkor Technology,Inc.,2Science Park Drive,Singapore 118222,Singapore
a r t i c l e i n f o Article history:
Received 22July 2009
Received in revised form 18January 2010Available online 24February 2010
a b s t r a c t
The increasing demand for portable electronics has led to the shrinking in size of electronic components and solder joint dimensions.The industry also made a transition towards the adoption of lead-free solder alloys,commonly based around the Sn–Ag–Cu alloys.As knowledge of the processes and operational reli-ability of these lead-free solder joints (used especially in advanced packages)is limited,it has become a major concern to characterise the mechanical performance of these interconnects amid the greater push for greener electronics by the European Union.
In this study,bulk solder tensile tests were performed to characterise the mechanical properties of SAC 105(Sn–1%wt Ag–0.5%wt Cu)and SAC 405(Sn–4%wt Ag–0.5%wt Cu)at strain rates ranging from 0.0088s à1to 57.0s à1.Solder joint array shear and tensile tests were also conducted on wafer-level chip scale package (WLCSP)specimens of different solder alloy materials under two test rates of 0.5mm/s (2.27s à1)and 5mm/s (22.73s à1).These WLCSP packages have an array of 12?12solder bumps (300l m in diameter);and double redistribution layers with a Ti/Cu/Ni/Au under-bump metallurgy (UBM)as their silicon-based interface structure.
The bulk solder tensile tests show that Sn–Ag–Cu alloys exhibit higher mechanical strength (yield stress and ultimate tensile strength)with increasing strain rate.A rate-dependent model of yield stress and ultimate tensile strength (UTS)was developed based on the test results.Good mechanical perfor-mance of package pull-tests at high strain rates is often correlated to a higher percentage of bulk solder failures than interface failures in solder joints.The solder joint array tests show that for higher test rates and Ag content,there are less bulk solder failures and more interface failures.Correspondingly,the aver-age solder joint strength,peak load and ductility also decrease under higher test rate and Ag content.The solder joint results relate closely to the higher rate sensitivity of SAC 405in gaining material strength which might prove detrimental to solder joint interfaces that are less rate sensitive.In addition,speci-mens under shear yielded more bulk solder failures,higher average solder joint strength and ductility than specimens under tension.
ó2010Elsevier Ltd.All rights reserved.
1.Introduction
Electronic components are shrinking in size to meet demands for lightweight and feature ?lled portable electronic products.This leads to decreasing solder joint dimensions,where mechanical reli-ability has become an issue [1],especially under high strain rate conditions during testing,transport and handling,impact loading under automotive [2]and consumer portable applications.
Tin lead alloy (SnPb)was commonly used as a solder material in microelectronic packaging,but it is also hazardous to the environ-ment and health.Therefore,the industry made a transition to lead-free solders,with the implementation a ban on lead (Pb)from elec-tronic products by the EU RoHS (restriction of the use of certain hazardous substances in electrical and electronic equipment)in
July 2006.The transition to lead-free solders is led by the widely adopted Sn–Ag–Cu (SAC)eutectic [3].However,some studies have shown that standard SAC alloys such as SAC 405(Sn–4%wt Ag–0.5%wt Cu)have poorer mechanical performance than eutectic SnPb under high strain rate conditions [4].Moreover,with the increasing popularity of portable devices,the performance of Sn–Ag–Cu solder joints under high strain rate and large rate ranges typical of drop impact situations is a major concern.
In this study,dogbone-shaped bulk material tensile tests were conducted to investigate the effect of strain rate and silver content on the material properties of Sn–Ag–Cu solders.Solder joint array shear and tensile experiments were conducted on WLCSP speci-mens of different alloy materials under different strain rates and loading orientations to investigate the effects of strain rate,silver content in Sn–Ag–Cu solder joints,and loading orientation on microelectronic packages.Failure analyses were also performed on the fractured dogbone-shaped bulk material test specimens and WLCSP solder joints.
0026-2714/$-see front matter ó2010Elsevier Ltd.All rights reserved.doi:10.1016/j.microrel.2010.01.043
*Corresponding author.
E-mail address:mpetanbc@https://www.doczj.com/doc/7717315386.html,.sg (V.B.C.Tan).
Microelectronics Reliability 50(2010)
564–576
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2.Experimental details and methodology
2.1.Test methodology for dogbone-shaped bulk solder tensile test Dogbone-shaped bulk solder specimens were used to perform tensile tests to characterise the mechanical properties of SAC 105and SAC 405.The bulk solder specimens had a gauge length of 19mm and diameter of 3mm as shown in Fig.1.The specimens were fabricated by machining from solder ingots and annealed at 70°C for 24h to reduce residual stresses.
The bulk solder tensile tests were conducted on a universal tes-ter.Three to four samples were tested at various test rates from 10mm/min (0.0088s à1)to 65,000mm/min (57.0s à1).2.2.Test methodology for solder joint array shear and tensile test Package pull and shear tests were conducted on wafer-level chip scale packages (WLCSP)with SAC 105or SAC 405alloy solder joints at 0.5mm/s and 5mm/s.Solder joint strength and ductility data are collected and failure analysis (FA)of the WLCSP joints performed via optical microscopy and SEM.The FA ?ndings are then statisti-cally tabulated and jointly analysed and correlated with test results.The WLCSP specimens were sawed out from board assemblies which were as-re?owed,unaged and non-solder mask de?ned,as shown in Fig.2.Each WLCSP specimen had an array of 12?12sol-der joints sandwiched between a die substrate and a printed circuit board.These WLCSP packages have solder bumps that are 300l m in diameter and double redistribution layers with a Ti/Cu/Ni/Au under-bump metallurgy (UBM)as their silicon-based interface structure.
The test samples were bonded onto ?xtures using a cyanoacry-late base adhesive and tested on an Instron Microtester.Specimens of two solder alloy materials (SAC 105and SAC 405)were tested at room conditions.Two to three samples were tested for each test parameter.The tests were carried out at test rates of 0.5mm/s (2.27s à1)and 5.0mm/s (22.73s à1)and two loading orientations of shear and tension,as shown in Fig.3.3.Results and discussion
3.1.Dogbone-shaped bulk solder tensile test results
The dogbone-shaped bulk solder specimens were tested un-der four test rates,with four samples per test rate.Nominal stress–strain data were derived from the load–displacement
raw data and plotted as shown in Fig.4a ,where the mechanical properties such as yield stress and ultimate tensile strength (UTS)were obtained.Yield stress was obtained at the point of 0.5%offset.True stress–strain curves were also plotted for comparison.
3.1.1.Effect of strain rate on bulk solder material properties
Fig.4b shows the true stress–strain curves of SAC 405,where the most representative sample curves of each of the four test rates were extracted and combined.It is observed that strain rate affects the bulk solder material properties,with increasing strength at higher rate of loading.Both Pang [5]and Che [6]had performed similar dogbone-shaped bulk solder tensile tests and also showed that material properties such as the yield stress and UTS of Sn–Ag–Cu solder alloys increase at higher strain rates.
3.1.1.1.Rate-dependent model development.Mechanical properties such as yield stress and UTS have linear logarithmic/power relationship with respect to strain rate.As such,a rate-dependent model can be developed to express this relationship quantitatively.
Nomenclature
Sn–Ag–Cu tin–silver–copper (Sn–Ag–Cu)alloy WLCSP wafer-level chip scale package r y yield stress,MPa r UTS ultimate tensile stress,MPa _e strain rate,s à1
Std.dev.standard deviation
DTBTSR ductile to brittle transition strain rate
SEM scanning electron microscope UBM under-bump metallization IMC inter-metallic compound Cu RDL copper redistribution layer
A T total failure cross sectional area,m 2
r ave
average solder joint array strength,MPa
F max
peak load from solder joint array shear and tensile tests,
N
Fig.1.Dimensions of dogbone bulk solder
specimen.Fig.2.Printed circuit board assembly and sawed WLCSP test
specimen.
Fig.3.Schematic of solder joint array shear and tensile tests.
Y.A.Su et al./Microelectronics Reliability 50(2010)564–576565
This is expressed in logarithmic scale as shown in Figs.5a and5b. Different solder alloy materials exhibit similar linear trend,with yield stress and UTS increasing at higher strain rate.The yield stress and UTS relationships with strain rate are obtained through curve?tting of the plots in Figs.5a and5b respectively to give:
r ye_eT
Sn—Ag—0:5Cu
?b1loge_eTtb2e1T
r
UTS
e_eTSn—Ag—0:5Cu?c1loge_eTtc2e2TFor ease of comparison with existing literature,the rate depen-dence in yield stress and UTS can be further expressed in the fol-lowing power relationships:r ye_eT
Sn—Ag—0:5Cu
?b3e_eTb4e3T
r
UTS
e_eTSn—Ag—0:5Cu?c3e_eTc4e4TThe coef?cients b1,b2,c1,c2,b3,b4,c3and c4are listed in Table1.
The material parameters of SAC105and Amkor’s internal data for SAC305are compared with the results obtained by Che[6], which are presented in Table2.
The development of such a rate-dependent model allows for better understanding of the relationships between mechanical properties and strain rate of applied loading.This enables the pre-diction of the mechanical properties at the strain rate of interest.It also enables the comparison of experimental results across similar literature.
3.1.1.2.Strain rate sensitivity.Referring to Figs.5a and5b,the high-er gradient of the linear curves translate to larger increase in strength with a given increase in strain rate,which can also be ex-pressed as higher strain rate sensitivity of material strength.The coef?cients b1and c1correspond to the gradient of the linear curves,which also directly relate to the strain rate sensitivity of yield stress and UTS respectively.Strain rate sensitivity can be use-ful in the study of the ductile to brittle transition strain rate (DTBTSR)in solder joints,which is affected by the sensitivity of bulk solder strength to strain rate,as depicted in Fig.12a.
3.1.2.Effect of Ag content on mechanical properties of bulk solder
Figs.5a,5b,6a and6b show the effect of silver(Ag)content on the yield stress and UTS of Sn–Ag–Cu solder alloy.Silver(Ag)con-tent of1%corresponds to SAC105and4%corresponds to SAC405. The yield stress and UTS of4%Ag content(SAC405)are consis-tently higher than1%Ag content(SAC105).Therefore,yield stress and UTS increase with higher Ag content in Sn–Ag–Cu alloys.
Che[6]had also obtained results that show higher yield stress, UTS and lower elongation(equivalent to ductility)in Sn–Ag–Cu al-loys with higher Ag content.Sn–Ag–Cu alloys have three phases of primary Sn,Ag3Sn and Cu6Sn5[4].Suh[4]had presented the mechanical properties of the three phases of Sn–Ag–Cu micro-structure and primary Sn has the lowest elastic modulus and strength of the three phases.Sn–Ag–Cu alloys are strengthened by the internal stress accumulated due to the difference in elastic modulus and volume fraction between Ag3Sn and Sn matrix[7]. Higher Ag content will increase the amount of Ag3Sn phase[2] and also lower the amount of primary Sn phase in Sn–Ag–Cu solder alloys[4].SAC105,with lower Ag content,is expected to have less Ag3Sn and more primary Sn phase.This contributes to the lower
566Y.A.Su et al./Microelectronics Reliability50(2010)564–576
strength of SAC 105compared to Sn–Ag–Cu alloys with higher Ag content [4].
3.1.3.Failure analysis of fractured dogbone-shaped bulk solder test specimens
Photographs of the fractured dogbone-shaped bulk solder test specimens were taken under an optical microscope.The side
views/overviews and closeup views of the specimens are presented in Fig.7.It is generally observed that 45°shearing is the primary mode of failure at low test/strain rates while necking is mostly ob-served at higher test/strain rates.At low strain rates,the solder al-loy is able to undergo slip deformation across the whole cross section of the dogbone shaped specimen,resulting in 45°shearing.As Sn–Ag–Cu alloy materials have high homologous temperature,there is not much strain hardening under low strain rate loading.At higher strain rates,the material is not able to respond to the ap-plied loading by slip deformation and experiences greater strain hardening,therefore necking occurs.From Fig.7,it can be observed that the diameter of fracture generally decreases at higher test/strain rate due to necking.
3.2.Solder joint array shear and tensile test results
3.2.1.Failure analysis of WLCSP joints
Optical microscopy and scanning electron microscopy (SEM)were used to examine the fractured solder joints,which are around 0.3mm in diameter each.Three modes of failure were identi?ed and categorised,namely:(1)bulk solder failure,(2)UBM IMC fail-ure and (3)pad matrix failure.Although there are other less pre-vailing failure modes observed (such as Cu RDL and Pad IMC),the three failure modes are chosen to be analysed due to their high frequency of occurrence (adding up to >90%of the failure modes of all the solder joints).They are also the most common failures iden-ti?ed by researchers [8,9].
Images of the respective failure modes at both the die/substrate side and board side were taken under the optical microscope and scanning electron microscope (SEM),as shown in Fig.8a–c .The photographs and SEM images are accompanied by cross sectional schematics to show the location of each failure mode.
3.2.1.1.Bulk solder failure.Bulk solder failure occurs by fracturing through the solder sphere as shown in Fig.8a.These failures are usually detected near the die side interface and can be easily iden-ti?ed by the silvery solder residue on the die substrate.From the board side,it can be recognised by the uneven fracture surface on the solder bump that is still attached to the board.
3.2.1.2.UBM IMC (under-bump metallization intermetallic compound)failure.Brittle interface failure occurs at the interfacial regions be-tween the bulk solder and the die substrate.The failure crack usu-ally occurs in the intermetallic compounds (IMC)or at the interfaces with the substrate.These intermetallic compounds are usually more brittle than the bulk solder [1].
UBM IMC failure refers to fracture at the intermetallic com-pound (IMC)interface layer between the UBM layer and the bulk solder.It can be identi?ed by a smooth grey surface on the die sub-strate.From the board side,it can also be identi?ed by its charac-teristic ring step and smooth surface at the top of the solder bump,shown in Fig.8b.
3.2.1.3.Pad matrix failure.Pad matrix failure is failure in the pad re-sin at the printed circuit board.It occurs at the matrix of the com-posite that makes up the board.It is identi?ed by the distinct red
Table 1
Material parameters of rate dependent yield stress and UTS of different solder alloys.
Logarithmic Power Yield stress UTS Yield stress UTS b 1
b 2
c 1c 2b 3b 4c 3c 4
SAC 105 2.92241.373 4.17454.71740.3780.077553.0930.081SAC 405
2.973
49.864
5.295
72.213
48.741
0.0596
69.633
0.0779
Table 2
Comparison of material parameters of rate dependent yield stress and UTS of different solder alloys between Amkor’s data and Che’s data based on Eqs.(3)and (4).
Yield stress UTS Amkor
Che Std.dev.Amkor Che Std.dev.b 3
b 4
SAC 10540.3832.15 5.8190.07750.05130.01853SAC 30552.74643.32 6.6650.07150.04450.01909c 3
c 4
SAC10553.09353.990.6340.0810.09180.00764SAC305
70.258
60.18
7.126
0.0777
0.0608
0.01195
Y.A.Su et al./Microelectronics Reliability 50(2010)564–576
567
copper pad on the solder bump that is attached to the die.It can also be identi?ed by ‘cratering’[3]in the board side,as shown in Fig.8c.Solder joints that were fractured during tests were examined to determine their modes of failure.Both Zhao [9]and Darveaux [8]had identi?ed three major modes of solder joint failures for
ball
Fig.7.Photographs of fractured dogbone-shaped bulk solder tensile test
specimens.
Fig.8.Schematic diagram of:(a)bulk solder failure,(b)UBM IMC failure and (c)pad matrix failure on die substrate,board,and solder joint cross section.
568Y.A.Su et al./Microelectronics Reliability 50(2010)564–576
grid array packages (BGAs),namely ductile bulk solder failure,brit-tle interface fracture at the intermetallic compound (IMC)layer and pad failure,shown in Fig.9.
Bulk solder failure is commonly regarded as a desirable mode of failure.A ductile bulk solder will deform when loading is applied,and this will minimise stresses in the interfacial regions due to bet-ter package/PCB curvature compliance.As failure in interfacial re-gions can occur under relatively small strain displacements and under rapid crack propagation rate,it will be preferable to reduce their occurrence to avoid catastrophic failures [8].Wong et al.noted that bulk solder failures correspond to high drop test life of ball grid array packages (BGAs),which is more desirable than IMC failures that correspond to low drop test life [10].Therefore,the observation of bulk solder failure is often related to better reli-ability of solder joints under high strain rate conditions.
3.2.2.Derivation of average solder joint array strength
The package pull/shear tests yielded load–displacement raw data,where the peak load was recorded.Darveaux had calculated the average stress in the solder joint array by dividing load over the sum of joint pad area [2].In this study,the WLCSP solder joint has varying cross sectional dimensions at the different regions of failure as shown in Fig.10.The number of solder joints that failed under each mode (n i )were counted based on photographs taken using an optical microscope,shown in Figs.11a and 11b .The fail-ure mode count (n i )was subsequently multiplied by the cross sec-tional area of that failure region in a single solder bump (a i ),to obtain the total affected cross sectional area for that particular mode of failure (A i ).The total failure cross sectional area (A T )can be obtained by summing up the failure cross sectional area of each failure mode (A i ).
Total failure cross sectional area,
A T ?
X
i
A i ?
X
i
a i n i e5T
where i =1corresponds to failure mode 1(e.g.bulk solder failure),i =2corresponds to failure mode 2(e.g.UBM IMC failure),etc.The average solder joint array strength (r ave )is calculated by dividing peak load (F max )over the total failure cross sectional area,A T .
Average solder joint array strength,
r ave ?
F max T
e6T
The peak loads obtained from the tests and the resulting peak stresses derived are given in Table 3.
As shown in the preceding paragraphs,there is a distinction be-tween bulk solder strength and solder joint strength.Bulk solder strength refers to the strength of the solder alloy in resisting fail-ure,whereas solder joint strength refers to the strength of the en-tire solder joint that consists of a multitude of different materials making up the joint,such as that of the bulk solder,PCB Cu pad,UBM,RDL Polyimide layer dielectric,and silicon die.Further,fail-ure of the solder joint may not lie within a particular material,but at the interfaces,e.g.UBM/RDL,solder/IMC,etc.Hence,a strong solder alloy may not necessarily create a strong and reliable solder joint.In general,solder joint reliability depends on its components’mechanical properties and interaction,and the proper selection of solder alloys and their properties on a particular joint structure and make-up would improve solder joint
life.
Fig.10.Typical cross sectional diameters of WLCSP solder
joint.
Fig.11a.Fractured solder joint array specimen after shear test (die
side).
Fig.11b.Fractured solder joint array specimen after shear test (pad
side).
Fig.9.Schematic of bulk solder failure and interface failures.
Y.A.Su et al./Microelectronics Reliability 50(2010)564–576569
3.2.3.Effect of strain rate
3.2.3.1.Ductile to brittle transition.Sn–Ag–Cu solder alloys gener-ally attain higher strength at higher strain rates [6,5].This is repre-sented in Fig.12a as an upward sloping line.At low strain rates,bulk solders have low strength,resulting in predominant failure in the bulk solders,depicted as ‘100%bulk solder failure’in Fig.12b .At higher strain rates,the bulk solders are stronger and more resistant to mechanical loading.The gain in strength also re-sulted in less deformation and compliance for the solder,and the applied loading stress will increasingly accumulate at the solder joint interfaces [8]due to their proximity to geometric discontinu-ities,resulting in more interfacial failures than bulk solder failures.As the test rates increases,there will eventually be no more bulk solder failures,being replaced entirely by failures occurring at the interface regions,depicted by ‘100%interface failure’in Fig.12b .As such,there is a transition from predominant ductile bulk solder failures to brittle interface failures with increasing strain rate,as shown in Fig.12b .
Darveaux had investigated the ductile to brittle transition strain rate (DTBTSR)by performing solder joint array tensile tests [2].The ductile to brittle transition strain rate (DTBTSR)was de?ned as the strain rate at which,50%of the joints fail at the pad interface’[2].In this study,‘ductile to brittle transition strain rate’will simply be the strain rate where 50%of the joints fail at the bulk solder,‘inter-facial failure’will collectively refer to all the other non bulk solder modes of failure.
It can be seen that the value of DTBTSR is a good gauge of what application strain rate the particular structure of solder joints can withstand before they start to fail abruptly at the joint interfaces.Knowledge of the DTBTSR values will allow researchers to design their packages to avoid excessive strain rate ranges that the joint would not be able to handle.
As it would be desirable to register a relatively high peak load and observe bulk solder failures in solder joints under high strain rate loading,from Fig.12b ,the preferred outcome then would be to design solder joints to have a higher ductile to brittle transition strain rate (DTBTSR)[8]for improved reliability under high strain rate conditions.
3.2.3.2.Effect of strain rate on solder joint array strength.Figs.13a and 13b show the average solder joint array strength of the tested specimens at two different test rates (0.5mm/s and 5.0mm/s)
un-
Fig.12a.Diagram of ductile to brittle transition in the relationship between solder joint strength and strain rate (adapted from [2]
).
Fig.12b.Diagram of ductile to brittle transition in the relationship between %bulk solder failures and strain rate (adapted from [2]).
Table 3
Solder joint array shear and tensile test results and derivation of average solder joint array strength.Solder alloy
Loading orientation
Test rate (mm/s)
Average solder joint array strength,r ave (MPa)
Failure mode count,n i Bulk solder
UBM IMC Cu RDL Pad matrix SAC 405
Shear 539.4401083330.543.663510107Tension
58.960001440.556.77060139SAC 105
Shear 536.890644770.554.79750070Tension
517.110511390.5
42.36
33
1
111
‘Pad IMC’mode of failure is not observed for SAC 105and SAC 405.
570Y.A.Su et al./Microelectronics Reliability 50(2010)564–576
der shear and tension.It is observed that the average solder joint array strength decreases at higher test/strain rates,regardless of solder alloy material and test orientation.Newman had observed higher strength at higher shear rates in solder ball impact shear tests [11],and Darveaux also obtained similar trend in solder joint array tensile tests [8].However,Darveaux also noted that solder joint strength declined at even higher strain rates when ‘interface failures start to occur’[8].From the experiments conducted in this study,interface failures had already started to occur in all the test specimens under the test rates of 0.5mm/s and 5mm/s.The test/strain rates employed in this study (0.5mm/s and 5mm/s)may al-ready be high enough to lead to the decline in average solder joint array strength.
Referring to Fig.12a of the ductile to brittle transition strain rate model,it can be seen that beyond the DTBTSR,there is a fall in interface strength of solder joints.Coupled with the increase in the amount of interface failures (see Fig.12b or Fig.17),this will lead to a decline in combined solder joint array strength at higher strain rate.The solder joint array test results in Figs.13a and 13b have also shown that average solder joint strength declined with higher test/strain rate.
3.2.3.3.Effect of strain rate on failure mode (under shear).(1)Bulk solder failure.
Fig.14a shows the percentage of bulk solder failures observed for the two different solder alloy solder joints tested under shear.At the higher test rate of 5mm/s,there is almost no more failure at the bulk solder,implying that all failures occur at the interfa-cial regions.As bulk solder becomes more load resistant and experience less deformation,it is less susceptible to mechanical failure at higher strain rates.In the transition from ductile to brittle failure,there is a decrease in bulk solder failures to even-tually no more bulk solder failure at higher strain rates [2],as shown in Fig.12b .This is also depicted in Fig.14b ,in which the experimental result of SAC 105from Fig.14a is superimposed onto Fig.12b of the ductile to brittle transition strain rate model.Fig.14b shows that the experimental results of SAC 105follow a similar trend as the ductile to brittle transition strain rate model.(2)UBM IMC failure.
Fig.14c shows the percentage of UBM IMC failures for the dif-ferent alloy solder joints under shear.Such failure modes are prominent at higher shear application rates due to the transition from ductile (cohesive)to brittle (interface)failures at increasingly higher strain rates.
Fig.14b.Experimental result of SAC 105(under shear)superimposed onto the ductile to brittle transition strain rate
model.
Y.A.Su et al./Microelectronics Reliability 50(2010)564–576571
(3)Pad matrix failure.
In Fig.14d,pad matrix failure is observed to be increasingly dominant for SAC105as the test rates increases for shear.In contrary,SAC405has a reduction of pad matrix failures at the higher test rate of5mm/s,as a result of a corresponding predom-inance of UBM IMC failures(see Fig.14c),which is also an interfa-cial mode of failure.
From the?gures for SAC405,the predominant failure mode changes from pad matrix failure at0.5mm/s to UBM IMC failure at5mm/s.(see Fig.14c).The observed failure modes transition may be a result of the UBM IMC layer being weaker than pad ma-trix layer at the higher test rate of5mm/s for SAC405.It is sur-mised that the higher rate sensitivity of SAC’s405mechanical strength may have contributed to the mode transition that is not yet observable for solder joints of the other two alloys.
Essentially,interfacial failures such as UBM IMC and pad matrix failures become more dominant at higher strain rates,due to the strengthening of the bulk solder.The?nal failure location depends on the components mechanical characteristics and is a competition among the different interfacial regions for failure as they share the applied load.
Figs.14e and14f shows the percentage of interface failures (UBM IMC and pad matrix failures)of SAC405and SAC105respec-tively at the two test rates under shearing.There is a higher per-centage of interface failures at the higher test rate.Non bulk solder modes of failures such as UBM IMC and pad matrix failures are interfacial failures that become more dominant at higher strain rate,due to the strengthening of the bulk solder.This is similar to Figs.12b and14b,depicting the transition from ductile to interface failures in solder joints at higher strain rates.
3.2.3.
4.Effect of strain rate on failure mode(under tension).(1)Bulk solder failure.
No bulk solder failure is observed for the two test rates.It is likely that the designated test rates of0.5mm/s and5mm/s have already surpassed the ductile to brittle transition strain rate (DTBTSR)for the three different alloyed joints and hence experi-ence100%interface failures shown in Fig.12b.In Fig.15a,the experimental result of SAC105is superimposed onto Fig.12b of the ductile to brittle transition strain rate model.Fig.15a also shows that the experimental results of SAC105correlates with the ductile to brittle transition strain rate model.
Fig.15a.Experimental result of SAC105(under tension)superimposed into the ductile to brittle transition strain rate model.
572Y.A.Su et al./Microelectronics Reliability50(2010)564–576
(2)UBM IMC failure.
Fig.15b shows the percentage of UBM IMC failures under ten-sion.Such failures decrease at higher test rates,which is in contrast to the results obtained from shear.As no bulk solder failure is ob-served under tension,the competing failure is between the various interfacial regions such as the UBM IMC and pad matrix.It can be seen that there is generally more pad matrix failures at the higher test rate of 5mm/s (see Fig.15c ),which results in correspondingly less UBM IMC failures.(3)Pad matrix failure.
A higher percentage of pad matrix failure is obtained at higher test rate for all alloys as shown in Fig.15c .The percentage of pad matrix failures increased to almost 100%at 5mm/https://www.doczj.com/doc/7717315386.html,parison with shear results imply indirectly that the pad matrix resin is weaker under tension.
Figs.15d and 15e show the percentage of interface failures (UBM IMC and pad matrix failures)of SAC 405and SAC 105respec-tively at the two test rates under tension.As there is no bulk solder failure at the two test rates,the solder joint array specimens expe-rienced 100%interface failures.From Fig.15a ,the two test rates of 0.5mm/s and 5mm/s may have already exceeded the DTBTSR of the SAC 105specimens under tension.
It was noted earlier that results from experiments showed a de-crease in average solder joint array strength with increasing strain rate.Further,a sharp decline in bulk solder failures and rise in interface failures are also observed at higher strain rates,that is similar to the trend in Fig.12b .Therefore,the combination of the above two conclusions imply that interfacial strength for the WLCSP joints are less rate-dependent than the alloys themselves and the related failure modes should be avoided especially under high strain rate/ranges conditions due to their likelihood of drastic failure as re?ected in the much lower solder joint array strength.3.2.4.Effect of Ag content
The dogbone (bulk solder)pull test results show evidently that bulk solder strength can be affected by Ag content in Sn–Ag–Cu solder alloys.The effect of Ag on bulk solder mechanical proper-ties had been discussed in Section 3.1.2‘Effect of Ag content on mechanical properties of bulk solder’,higher Ag content in Sn–Ag–Cu solder alloys generally leads to higher bulk solder strength.From the materials perspective,the trace amount of silver inclu-
Y.A.Su et al./Microelectronics Reliability 50(2010)564–576573
sions help impede the movement of dislocations across the alloy grains to result in the strengthening of the bulk alloy against deformation.Silver(Ag),being a harder substance than tin(Sn) and copper(Cu)also hardens the overall bulk alloy.
574Y.A.Su et al./Microelectronics Reliability50(2010)564–576
3.2.
4.1.Effect of Ag content on failure mode.Figs.16a and16b show the distribution of the various modes of failure under shear.
In Fig.16a,SAC105experiences more bulk solder failures than SAC405under0.5mm/s shear.It is expected that higher Ag con-tent will result in higher bulk alloy strength under the same test rate,therefore,SAC105(1%Ag)joints are seen to obtain lower strength than SAC405(4%Ag)at0.5mm/s.For5mm/s,the trend is contradictory as the higher rate has already caused the increase in interface failures for SAC405that resulted in a decline in overall joint strength.
In Fig.16b,SAC405has higher percentage of UBM IMC failures than SAC105under5mm/s shear.As SAC405has higher strength than SAC105,SAC405bulk solder is more load resistant,thereby transferring more loading stress to the interfaces,resulting in more interface failures than in SAC105solder joints.Kim et al.con-ducted drop tests and found that IMC cracking at the package side was predominant in SAC405,while SAC105had more bulk solder failures[12].Sn–Ag–Cu solder joint arrays with lower Ag content will yield more bulk solder failures.
The tensile test results have predominant pad matrix failure and no bulk solder failure,shown in Figs.16c and16d.
3.2.
4.2.Effect of Ag content on solder joint array strength.From Figs. 13a and13b,there seems to be no obvious trend in the average sol-der joint array strength under the effect of Ag content in Sn–Ag–Cu solders.However,analysing the peak loads,where the average sol-der joint array strength was derived,display a general trend. Fig.16e shows that the peak load of SAC105(1%Ag)is generally higher than the peak load of SAC405(4%Ag).Solder joint array specimens with low Ag content will require a higher peak load to fail as compared to their higher Ag content counterparts,which translates to better mechanical performance of low Ag packages under loading.The discrepancy in the values and trends between the peak load and solder joint array strength is a result of the dif-ferences in the total failure area due the varying modes of failure, as described in the Section3.2.2.
Solder joint arrays with lower Ag content are able to withstand higher peak loads before failure occurs.Therefore,solder joints with Sn–Ag–Cu alloys of low Ag content will improve the reliability of microelectronic packages under high strain rate conditions.
3.2.5.Effect of loading orientation
The experimental results in Fig15a have shown that there is no bulk solder failure at the two test rates under tension.It is likely that the test/strain rates that are used in the solder joint ar-ray experiments are already higher than their ductile to brittle transition strain rate(DTBTSR).A likely scenario of the SAC105 solder joints under both shear and tension is depicted in the fol-lowing example based on certain assumptions and experimental results.
The results of SAC105under different orientations will be used to illustrate the effect of loading orientation on the ductile to brittle transition strain rate of its solder joint in Fig.17.The characteristics of SAC105mentioned above are applied onto Fig.12b to derive Fig.17.The shear tests on SAC105specimens yielded52%bulk solder failure at0.5mm/s.However,the tensile tests yielded no bulk solder failure at the same test rate.It is as-sumed that under both tension and shear loading,the specimens have similar strain rate sensitivity(i.e.,same gradient in the tran-sition lines in Fig.17)from0%to100%interface failures.In Fig.17,the DTBTSR corresponds to the point marked by a black dot.It can be seen that SAC105specimens under tension have lower ductile to brittle transition strain rate(DTBTSR)than spec-imens under shear.
The specimens of SAC405also yielded some bulk solder failures under shear but none under tension.Therefore similar results as in the case of SAC105presented above can also be derived,to reach the same conclusions that specimens under tension generally have lower ductile to brittle transition strain rate(DTBTSR)than speci-mens under shear.This is similar to the results obtained by Darve-aux,where the ductile to brittle transition strain rate(DTBTSR) under tensile loading is lower than DTBTSR under shear loading [2].
In Section3.2.3‘Effect of strain rate:Ductile to brittle transition’, it was mentioned that it is desirable for solder joints to have higher DTBTSR,coupled with higher joint strength,for improved solder joint reliability.Therefore,it is preferable for solder joints to expe-rience some amount of shear under high strain rate loading condi-tions such as drop impact.
Fig.18shows the effect of loading orientation on the load dis-placement pro?les of SAC105specimens at a test rate of 0.5mm/s.The specimen under shear has higher peak load and duc-tility(larger displacement)and requires a greater work done than a similar specimen under tension.
Figs.13a and13b have shown that the tested specimens under 5mm/s shear have higher average solder joint array strength than those under5mm/s tension.The results illustrate the higher mechanical performance of microelectronic packages under shear loading compared to tensile loading,neglecting the effect of fatigue resistance(crack propagation propensity)through the bulk or interfacial material.The study shows that it would be preferable to design WLCSP or microelectronic packages to experience some
Fig.17.Illustration of lower ductile to brittle transition strain rate(DTBTSR)of SAC 105test specimens under different loading orientations.
Y.A.Su et al./Microelectronics Reliability50(2010)564–576575
amount of shear,than complete tensile stresses to curb solder-joint damage induced during shock/drop impact.
4.Conclusions
The mechanical properties of Sn–Ag–Cu lead-free solder alloy materials were investigated in two directions of study.Firstly,bulk solder tensile tests results were used to investigate the mechanical properties of bulk solder alloy materials and secondly,solder joint array shear and tensile tests were performed to investigate the per-formance of solder joints of different solder alloy materials.
From the results of bulk solder tensile tests,the bulk material properties of SAC105and SAC405were derived and a rate-depen-dent model of yield stress and UTS was developed.The following conclusions were derived from the analysis of the experimental results:
(1)Solder alloy materials such as SAC105and SAC405exhibit
higher mechanical strength(yield stress and UTS)with increasing strain rate.
(2)Rate-dependent properties such as mechanical strength
exhibit linear relationship with respect to strain rate expressed in logarithmic/power scale.Material parameters obtained from these rate dependent relationships can be compared across literature.
(3)Bulk solder material properties such as strength and ductil-
ity vary with higher Ag content in Sn–Ag–Cu alloys.
In the second direction of study,solder joint array shear and tensile tests were performed using WLCSP with different solder al-loy materials(SAC105,SAC405)under different test rates of 0.5mm/s(2.27sà1)and5mm/s(22.73sà1).The following conclu-sions were derived from the analysis of the solder joint array test results:
(1)Average solder joint array strength decreases at higher strain
rates.
(2)For similar alloys,more interface failures are observed at
higher strain rate.
(3)High Ag content Sn–Ag–Cu joints generally yield more inter-
face failures,and lower peak loads than those with lower Ag content.
(4)Solder joints under shear loading produced more bulk solder
failures and higher average solder joint strength,peak load and ductility than those under tensile loading.
Having more bulk solder failures and higher average solder joint array strength are some of the desirable characteristics commonly associated with better mechanical reliability of solder joints in microelectronic packages.The results of this paper are in line with much literature works[2,8,12]and show that microelectronic packages with low Ag content,such as SAC105solder joints,gen-erally exhibit much of these desirable characteristics compared to SAC405,at least for the test variables and package types tested in this paper.
Acknowledgements
The authors would like to thank Alvin Goh and Joe Low for facil-ities support at the NUS Impact Mechanics Laboratory and Mr. Shane Loo for his support and assistance in the project. References
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— 1 — 焊点可靠性之焊点寿命预测 在产品设计阶段对SMT 焊点的可能服役期限进行预测，是各大电子产品公司为保证电子整机的可靠性所必须进行的工作，为此提出了多种焊点寿命预测模型。 (1) 基于Manson-Coffin 方程的寿命预测模型 M-C 方程是用于预测金属材料低周疲劳失效寿命的经典经验方程[9]。其基本形式如下： C N p f =ε?β (1-1) 式中 N f — 失效循环数； ?εp — 循环塑性应变范围； β, C — 经验常数。 IBM 的Norris 和Landzberg 最早提出了用于软钎焊焊点热疲劳寿命预测的M-C 方程修正形式[2]： )/exp()(max /1kT Q Cf N n p m f -ε?= (1-2) 式中 C, m, n — 材料常数； Q — 激活能； f — 循环频率； k — Boltzmann 常数； T max — 温度循环的最高温度。 Bell 实验室的Engelmaier 针对LCCC 封装SMT 焊点的热疲劳寿命预测对M-C 方程进行了修正[10]： c f f N /1'221???? ??εγ?= (1-3) )1ln(1074.1106442.024f T c s +?+?--=-- (1-4) 式中 ?γ — 循环剪切应变范围； f 'ε— 疲劳韧性系数，2f 'ε=0.65； c — 疲劳韧性指数； T s — 温度循环的平均温度。 采用M-C 型疲劳寿命预测方程，关键在于循环塑性应变范围的确定。主要有两种方法：一种是解析法[10,11]，通过对焊点结构的力学解析分析计算出焊点在热循环过程中承受的循环应变范围，如Engelmaier 给出[10]：
金属材料焊接性知识要点 1. 金属焊接性：指同质材料或异质材料在制造工艺条件下，能够形成完整接头并满足预期使用要求的能力。包括（工艺焊接性和使用焊接性）。 2. 工艺焊接性：金属或材料在一定的焊接工艺条件下，能否获得优质致密无缺陷和具有一定使用性能的焊接接头能力。 3. 使用焊接性：指焊接接头和整体焊接结构满足各种性能的程度，包括常规的力学性能。 4. 影响金属焊接性的因素：1、材料本因素2、设计因素3、工艺因素4、服役环境 5. 评定焊接性的原则:（1）评定焊接接头中产生工艺缺陷的倾向，为制定合理的焊接工艺提供依据；（2）评定焊接接头能否满足结构使用性能的要求。 6. 实验方法应满足的原则：1可比性 2针对性 3再现性 4经济性 7. 常用焊接性试验方法： A:斜Y坡口焊接裂纹试验法: 此法主要用于评定碳钢和低合金高强钢焊接热影响区对冷裂纹的敏感性。 B:插销试验 C:压板对接焊接裂纹试验法 D:可调拘束裂纹试验法 一问答：1、“小铁研”实验的目的是什么，适用于什么场合？了解其主要实验步骤，分析影响实验结果稳定性的因素有哪些？ 答：1、目的是用于评定用于评定碳钢和低合金高强钢焊接热影响区对冷裂纹的敏感性。评定碳钢和低合金高强钢焊接热影响区对冷裂纹的敏感性时，影响结果稳定因素焊接接头拘束度预热温度角变形和未焊透。（一般认为低合金钢“小铁研实验”表面裂纹率小于20%时。用于一般焊接结构是安全的） 2、影响工艺焊接性的主要因素有哪些？ 答：影响因素：（1）材料因素包括母材本身和使用的焊接材料，如焊条电弧焊的焊条、埋弧焊时的焊丝和焊剂、气体保护焊时的焊丝和保护气体等。 （2）设计因素焊接接头的结构设计会影响应力状态，从而对焊接性产生影响。 （3）工艺因素对于同一种母材，采用不同的焊接方法和工艺措施，所表现出来的焊接性有很大的差异。 （4）服役环境焊接结构的服役环境多种多样，如工作温度高低、工作介质种类、载荷性质等都属于使用条件。 3、举例说明有时工艺焊接性好的金属材料使用焊接性不一定好。 答：金属材料使用焊接性能是指焊接接头或整体焊接结构满足技术条件所规定的各种使用性能主要包括常规的力学性能或特定工作条件下的使用性能，如低温韧性、断裂韧性、高温蠕变强度、持久强度、疲劳性能以及耐蚀性、耐磨性等。而工艺焊接性是指金属或材料在一定的焊接工艺条件下，能否获得优质致密、无缺陷和具有一定使用性能的焊接接头的能力。比如低碳钢焊接性好，但其强度、硬度却没有高碳钢好。 4、为什么可以用热影响区最高硬度来评价钢铁材料的焊接冷裂纹敏感性？焊接工艺条件对热影响区最高硬度有什么影响？ 答：因为（1）.冷裂纹主要产生在热影响区； （2）其直接评定的是冷裂纹产生三要素中最重要的，接头淬硬组织，所以可以近似用来评价冷裂纹。 一般来说，焊接接头包括热影响区，它的硬度值相对于母材硬度值越高，证明焊接接头的
SMT焊点可靠性研究 前言 近几年﹐随着支配电子产品飞速发展的高新型微电子组装技术--表面组装技术(SMT)的飞速发展﹐SMT焊点可靠性问题成为普遍关注的焦点问题。 与通孔组装技术THT(Through Hole Technology)相比﹐SMT在焊点结构特征上存在着很大的差异。THT焊点因为镀通孔内引线和导体铅焊后﹐填缝铅料为焊点提供了主要的机械强度和可靠性﹐镀通孔外缘的铅焊圆角形态不是影响焊点可靠性的主要因素﹐一般只需具有润湿良好的特征就可以被接受。但在表面组装技术中﹐铅料的填缝尺寸相对较小﹐铅料的圆角(或称边堡)部分在焊点的电气和机械连接中起主要作用﹐焊点的可靠性与THT焊点相比要低得多﹐铅料圆角的凹凸形态将对焊点的可靠性产生重要影响。 另外﹐表面组装技术中大尺寸组件(如陶瓷芯片载体)与印制线路板的热膨胀系数相差较大﹐当温度升高时﹐这种热膨胀差必须全部由焊点来吸收。如果温度超过铅料的使用温度范围﹐则在焊点处会产生很大的应力最终导致产品失效。对于小尺寸组件﹐虽然因材料的CTE 失配而引起的焊点应力水平较低﹐但由于SnPb铅料在热循环条件下的粘性行为(蠕变和应力松弛)存在着蠕变损伤失效。因此﹐焊点可靠性问题尤其是焊点的热循环失效问题是表面组装技术中丞待解决的重大课题。 80年代以来﹐随着电子产品集成水平的提高,各种形式﹑各种尺寸的电子封装器件不断推出﹐使得电子封装产品在设计﹑生产过程中,面临如何合理地选择焊盘图形﹑焊点铅料量以及如何保证焊点质量等问题。同时﹐迅速变化的市场需求要求封装工艺的设计者们能快速对新产品的性能做出判断﹑对工艺参数的设置做出决策。目前﹐在表面组装组件的封装和引线设计﹑焊盘图形设计﹑焊点铅料量的选择﹑焊点形态评定等方面尚未能形成合理统一的标准或规则﹐对工艺参数的选择﹑焊点性能的评价局限于通过大量的实验估测。因此﹐迫切需要寻找一条方便有效的分析焊点可靠性的途径﹐有效地提高表面组装技术的设计﹑工艺水平。 研究表明﹐改善焊点形态是提高SMT焊点可靠性的重要途径。90年代以来﹐关于焊点形成及焊点可靠性分析理论有大量文献报导。然而﹐这些研究工作都是专业学者们针对焊点
计算机系统的焊点可靠性试验(doc 5页)
焊点可靠性试验的计算机模拟 本文介绍，与实际的温度循环试验相比，计算机模拟提供速度与成本节约。 在微电子工业中，一个封装的可靠性一般是通过其焊点的完整性来评估的。锡铅共晶与近共晶焊锡合金是在电子封装中最常用的接合材料，提供电气与温度的互联，以及机械的支持。由于元件内部散热和环境温度的变化而产生的温度波动，加上焊锡与封装材料之间热膨胀系统(CTE)的不匹配，造成焊接点的热机疲劳。不断的损坏最终导致元件的失效。 在工业中，决定失效循环次数的标准方法是在一个温室内进行高度加速的应力试验。温度循环过程是昂贵和费时的，但是计算机模拟是这些问题的很好的替代方案。模拟可能对新的封装设计甚至更为有利，因为原型试验载体的制造成本非常高。本文的目的是要显示，通过在一个商业有限单元(finite element)代码中使用一种新的插入式专门用途的材料子程序，试验可以在计算机屏幕上模拟。 建模与试验 宁可通过计算程序试验来决定焊点可靠性的其中一个理由是缺乏已验证的专用材料模型和软件包。例如，市场上现有的所有主要的商业有限单元分析代码都对应力分析有效，但是都缺乏对焊点以统一的方式进行循环失效分析的能力。该过程要求一个基于损伤机制理论的专门材料模型和在实际焊点水平上的验证。可以肯定的是，所有主要的有限单元分析代码都允许用户实施其自己的用户定义的插入式材料子程序。 直到现在，还不可能测量疲劳试验期间在焊点内的应力场，这对确认材料模型是必须的。在Buffalo大学的电子封装实验室(UB-EPL)开发的一个Moiré 干涉测量系统允许在疲劳试验到失效期间的应力场测试。 基于热力学原理的疲劳寿命预测模型也已经在UB-EPL开发出来，并用于实际的BGA封装可靠性试验的计算机模拟。在焊点内的损伤，相当于在循环热机负载下材料的退化，用一个热力学构架来量化。损伤，作为一个内部状态变量，结合一个基于懦变的构造模型，用于描述焊点的反映。该模型通过其用户定义的子程序实施到一个商业有限单元包中。 预测焊点的可靠性 焊接点的疲劳寿命预测对电子封装的可靠性评估是关键的。在微电子工业中预测失效循环次数的标准方法是基于使用通过试验得出的经验关系式。如果
焊点可靠性试验的计算机模拟 本文介绍，与实际的温度循环试验相比，计算机模拟提供速度与成本节约。 在微电子工业中，一个封装的可靠性一般是通过其焊点的完整性来评估的。锡铅共晶与近共晶焊锡合金是在电子封装中最常用的接合材料，提供电气与温度的互联，以及机械的支持。由于元件内部散热和环境温度的变化而产生的温度波动，加上焊锡与封装材料之间热膨胀系统(CTE)的不匹配，造成焊接点的热机疲劳。不断的损坏最终导致元件的失效。 在工业中，决定失效循环次数的标准方法是在一个温室内进行高度加速的应力试验。温度循环过程是昂贵和费时的，但是计算机模拟是这些问题的很好的替代方案。模拟可能对新的封装设计甚至更为有利，因为原型试验载体的制造成本非常高。本文的目的是要显示，通过在一个商业有限单元(finite element)代码中使用一种新的插入式专门用途的材料子程序，试验可以在计算机屏幕上模拟。建模与试验 宁可通过计算程序试验来决定焊点可靠性的其中一个理由是缺乏已验证的专用材料模型和软件包。例如，市场上现有的所有主要的商业有限单元分析代码都对应力分析有效，但是都缺乏对焊点以统一的方式进行循环失效分析的能力。该过程要求一个基于损伤机制理论的专门材料模型和在实际焊点水平上的验证。可以肯定的是，所有主要的有限单元分析代码都允许用户实施其自己的用户定义的插入式材料子程序。 直到现在，还不可能测量疲劳试验期间在焊点内的应力场，这对确认材料模型是必须的。在Buffalo大学的电子封装实验室(UB-EPL)开发的一个Moiré干涉测量系统允许在疲劳试验到失效期间的应力场测试。 基于热力学原理的疲劳寿命预测模型也已经在UB-EPL开发出来，并用于实际的BGA封装可靠性试验的计算机模拟。在焊点内的损伤，相当于在循环热机负载下材料的退化，用一个热力学构架来量化。损伤，作为一个内部状态变量，结合一个基于懦变的构造模型，用于描述焊点的反映。该模型通过其用户定义的子程序实施到一个商业有限单元包中。 预测焊点的可靠性 焊接点的疲劳寿命预测对电子封装的可靠性评估是关键的。在微电子工业中预测失效循环次数的标准方法是基于使用通过试验得出的经验关系式。如果使用一个分析方法，通过都是使用诸如Coffin-Manson(C-M)这样的经验曲线。通常，
金属焊接性与焊接结构设计 专业_________班级_______学号_______姓名___________ 10-1 判断题（正确的画O，错误的画×） 1．金属的焊接性不是一成不变的。同一种金属材料，采用不同的焊接方法及焊接材料，其焊接性可能有很大差别。（O）2．焊接中碳钢时，常采用预热工艺。预热对减小焊接应力十分有效。同时，预热也可防止在接头上产生淬硬组织。（O）3．根据等强度原则，手工电弧焊焊接400MPa级的15MnV钢，需使用结426和结427（或结422、结423）焊条。（×） 10-2 选择题 1．不同金属材料的焊接性是不同的。下列铁碳合金中，焊接性最好的是（D）。 A．灰口铸铁；B．可锻铸铁；C．球墨铸铁； D．低碳钢；E．中碳钢；F．高碳钢。 2．焊接梁结构，焊缝位置如图10-l所示，结构材料为16Mn钢，单件生产。上、下翼板的拼接焊缝A应用（B）方法和（E）焊接材料；翼板和腹板的四条长焊缝B 宜采用（A）方法焊接，使用的焊接材料为（H）；筋板焊缝C应采用（A）方法焊接，焊接材料为（E）。 A．埋弧自动焊；B．手工电弧焊；C．氩弧焊；D．电渣焊；E．结507； F．结422；G．结427；H．H08MnA和焊剂431；I．H08MnSiA； J．H08A和焊剂130。 图10-1 10-3 填空题
1．图10-2所示为汽车传动轴，由锻件45钢和钢管Q235钢焊接而成。大批量生产时，合适的焊接方法为（CO 气体保护焊）；使用的焊接材料为（H08MnSiA）。 2 图10-2 图10-3 2．汽车车轮由轮圈和辐板组成，材料均为Q235钢，如图10-3所示。大批量生产时，轮圈由卷板机卷成，再经（闪光对焊）焊接而成；而轮圈与辐板则用（CO 气体保护 2 焊）焊接连为一体，焊接材料为（H08MnSiA）。 10-4 应用题 1．在长春地区用30mm厚的16Mn钢板焊接一直径为20m的容器。16Mn的化学成分如下：C＝0.12～0.20％；Si＝0.20～0.55％；Mn＝1.20～1.60％；P、S＜0.045％。 （1）计算16Mn的碳当量； （2）判断16Mn的焊接性； （3）夏季施工时是否需要预热?冬季施工时是否需要预热？如需预热，预热温度应为多少? =0.16+1.40/6 =0.39 16Mn的焊接性良好 夏季施工时不需要预热。冬季施工时需要预热，预热温度应100-150℃。 2．修改焊接结构的设计(焊接方法不变) （1）钢板的拼焊（电弧焊），如图10-4。
金属熔焊原理考点 一．基础题： 1 焊接参数包括：焊接电流、电弧电压、焊接速度、线能量等。（参照课本P15图1-6） 2 焊条的平均熔化速度、熔敷速度均与电流成正比。 3 短路过渡的熔滴质量和过渡周期主要取决于电弧长（电弧电压），随电弧长度的增加，熔滴质量与过渡周期增大。当电弧长度到达一定值时，熔滴质量与过渡周期突然增大，这说明熔滴的过渡形式发生了变化，如果电弧长度不变，增大电流则过渡频率增高，熔滴变细。 4 一般情况下，增大焊接电流，熔宽减小，熔深增大；增大电弧电压，熔宽增大，熔深减小。 5 熔池的温度分布极其不均匀（熔池中部温度最高）。 6 焊接方法的保护方式：手弧焊（气-渣联合保护），埋弧焊、电渣焊（熔渣保护），氩弧焊 CO2焊、等离子焊（气体保护）。 7 焊接化学冶金过程是分区域连续进行的。 8 焊接化学冶金反应区：手工焊有药皮反应区、熔滴反应区、熔池反应区三个反应区；熔化极气保焊只有熔滴和熔池两个反应区；不填充金属的气焊、钨极氩弧焊和电子束焊只有熔池反应区。 9 熔滴阶段的反应时间随焊接电流的增加而变短，随电弧电压的增加而变长。 10 焊接材料只影响焊缝成分而不影响热影响区。 11 焊接区周围的空气是气相中氮的主要来源。 12 熔渣在焊接过程中的作用：机械保护、改善焊接工艺性能、冶金处理。 13 分子理论中酸碱性以1为界点，原子理论中，以0为界点。 14影响FeO分配系数的主要因素有：温度和熔渣的性质。 15焊缝金属的脱氧方式：先期脱氧、沉淀脱氧、扩散脱氧。 16脱硫比脱磷更困难。 17随焊芯中碳含量的增加，焊接时不仅焊缝中的气孔、裂纹倾向增大，并伴有较大飞溅，是焊接稳定性下降。 18焊条的冶金性能是指其脱氧、去氢、脱硫磷、掺合金、抗气孔及抗裂纹的能力，最终反映在焊缝金属的化学成分、力学性能和焊接缺陷的形成等方面。 19 焊剂按制造方法分为：熔炼焊剂和非熔炼焊剂。 20 焊丝的分类：实芯焊丝和药芯焊丝。 21 焊接中的偏析形式：显微偏析、区域偏析、层状偏析。 22 相变组织（二次结晶组织）主要取决于焊缝化学成分和冷却条件。 23焊接热循环的基本参数：加热速度、最高加热速度、相变温度以上停留的时间、冷却速度或冷却时间t8/5、t8/3、t100。 24 产生冷裂纹的三要素：拘束应力、淬硬组织、氢的作用 25冷裂纹的断口组织，宏观上看冷裂纹的断口具有淬硬性断裂的特征，表面有金属光泽，呈人字形发展，从微观上看，裂纹多起源于粗大奥氏体晶粒的晶界交错处。 26 冷裂纹的种类：延迟裂纹、淬硬脆化裂纹、低塑性脆化裂纹。 27 熔滴过度的作用力：重力、表面张力、电磁压缩力及电弧吹力等。 28活性熔渣对焊缝金属的氧化形式：扩散氧化、置换氧化。 29 熔合比影响焊缝的化学成分、金属组织和机械性能。局部熔化的母材将对焊缝的成分起到稀释作用。 30 焊接过程中对金属的保护有气保护、气-渣联合保护、渣保护、自保护。 二．名词解释： 1 焊接温度场：焊接过程中某一瞬时间焊接接头上个点的温度分布状态。 2 焊缝金属的熔合比：熔化焊时，被熔化的母材在焊缝金属中所占的百分比。 3 药皮重量系数：单位长度药皮与焊芯的质量比。 4 随温度降低黏度缓慢增加的称为长渣。随温度降低黏度迅速降低的称为短渣。 5 合金元素的过度系数：指某合金元素在熔敷金属中的实际质量分数与其在焊材中的原始质量分数之
焊接原理与焊锡性 收藏此信息打印该信息添加：用户投稿来源：未知 1、Abietic Acid松脂酸 是天然松香（Rosin）的主要成份，占其重量比的34％。在焊接的高温下，此酸能将铜面的轻微氧化物或钝化物予以清除，使得清洁铜面可与熔锡产生"接口合金共化"（IMC）而完成焊接。此松脂酸在常温中很安定，不会腐蚀金属。 2、Angle of Contack 接触角 广义是指液体落在固体表面时，其边缘与固体外表在截面上所形益的夹角。在PCB 的狭义上是指焊锡与铜面所形成的Θ角，又称之为双反斜角 （Dihedrel Angle）或直接称为 Contact Angle。 3、Blow Hole 吹孔 指完工的 PTH 铜壁上，可能有破洞(Void 俗称窟窿)存在。当板子在下游进行焊锡时，可能会造成破洞中的残液在高温中迅速气化而产生压力，往外向孔中灌入的熔锡吹出。冷却后孔中之锡柱会出现空洞。这种会吹气的劣质 PTH，特称为"吹孔"。吹孔为 PCB 制程不良的表征，必须彻底避免才能在业界立足。 4、Brazing 硬焊 是指采用含银的铜锌合金焊条，其焊温在425~870℃下进行熔接(Welding)方式，比一般电子工业常见软焊或焊(Soldering)，在温度及强度方面都比较高。 5、Cold Solder Joint 冷焊点 焊锡与铜面间在高温焊接过程中，必须先出现 Cn Sn 的"接口合金共化物"(IMC)层，才会出现良好的沾锡或焊锡性(Solderability)。当铜面不洁、热量不足，或焊锡中杂质太多时，都无法形成必须的 IMC(Eta Phase)，将出现灰暗多凹坑不平。且结构强度也不足的焊点，系由焊锡冷凝而形成，但未真正焊牢的焊点，特称为"冷焊点"，或俗称冷焊。 6、Contact Angle 接触角 一般泛指液体与固接触时，其交界边缘，在液体与固体外表截面上，所呈现的交接角度，谓之 Contact Angle。 7、Dewetting 缩锡 指高温熔融的焊锡与被焊物表面接触及沾锡后，当其冷却固化即完成焊接作用得到焊点(Solder Joint)。正常的焊点或焊面，其已固化的锡面都应呈现光泽平滑的外观，是为焊锡性(Solderability)良好的表征。所谓 Dewetting 是指焊点或焊面呈高低不平、多处下陷，或焊锡面支离破碎甚至曝露底金属，或焊点外缘无法顺利延伸展开，截面之接触角大于 90 度者，皆称为"缩锡"。其基本原因是底
Rate-dependent properties of Sn–Ag–Cu based lead-free solder joints for WLCSP Y.A.Su a ,L.B.Tan a ,T.Y.Tee b ,V.B.C.Tan a,* a National University of Singapore,Department of Mechanical Engineering,9Engineering Drive 1,Singapore 117576,Singapore b Amkor Technology,Inc.,2Science Park Drive,Singapore 118222,Singapore a r t i c l e i n f o Article history: Received 22July 2009 Received in revised form 18January 2010Available online 24February 2010 a b s t r a c t The increasing demand for portable electronics has led to the shrinking in size of electronic components and solder joint dimensions.The industry also made a transition towards the adoption of lead-free solder alloys,commonly based around the Sn–Ag–Cu alloys.As knowledge of the processes and operational reli-ability of these lead-free solder joints (used especially in advanced packages)is limited,it has become a major concern to characterise the mechanical performance of these interconnects amid the greater push for greener electronics by the European Union. In this study,bulk solder tensile tests were performed to characterise the mechanical properties of SAC 105(Sn–1%wt Ag–0.5%wt Cu)and SAC 405(Sn–4%wt Ag–0.5%wt Cu)at strain rates ranging from 0.0088s à1to 57.0s à1.Solder joint array shear and tensile tests were also conducted on wafer-level chip scale package (WLCSP)specimens of different solder alloy materials under two test rates of 0.5mm/s (2.27s à1)and 5mm/s (22.73s à1).These WLCSP packages have an array of 12?12solder bumps (300l m in diameter);and double redistribution layers with a Ti/Cu/Ni/Au under-bump metallurgy (UBM)as their silicon-based interface structure. The bulk solder tensile tests show that Sn–Ag–Cu alloys exhibit higher mechanical strength (yield stress and ultimate tensile strength)with increasing strain rate.A rate-dependent model of yield stress and ultimate tensile strength (UTS)was developed based on the test results.Good mechanical perfor-mance of package pull-tests at high strain rates is often correlated to a higher percentage of bulk solder failures than interface failures in solder joints.The solder joint array tests show that for higher test rates and Ag content,there are less bulk solder failures and more interface failures.Correspondingly,the aver-age solder joint strength,peak load and ductility also decrease under higher test rate and Ag content.The solder joint results relate closely to the higher rate sensitivity of SAC 405in gaining material strength which might prove detrimental to solder joint interfaces that are less rate sensitive.In addition,speci-mens under shear yielded more bulk solder failures,higher average solder joint strength and ductility than specimens under tension. ó2010Elsevier Ltd.All rights reserved. 1.Introduction Electronic components are shrinking in size to meet demands for lightweight and feature ?lled portable electronic products.This leads to decreasing solder joint dimensions,where mechanical reli-ability has become an issue [1],especially under high strain rate conditions during testing,transport and handling,impact loading under automotive [2]and consumer portable applications. Tin lead alloy (SnPb)was commonly used as a solder material in microelectronic packaging,but it is also hazardous to the environ-ment and health.Therefore,the industry made a transition to lead-free solders,with the implementation a ban on lead (Pb)from elec-tronic products by the EU RoHS (restriction of the use of certain hazardous substances in electrical and electronic equipment)in July 2006.The transition to lead-free solders is led by the widely adopted Sn–Ag–Cu (SAC)eutectic [3].However,some studies have shown that standard SAC alloys such as SAC 405(Sn–4%wt Ag–0.5%wt Cu)have poorer mechanical performance than eutectic SnPb under high strain rate conditions [4].Moreover,with the increasing popularity of portable devices,the performance of Sn–Ag–Cu solder joints under high strain rate and large rate ranges typical of drop impact situations is a major concern. In this study,dogbone-shaped bulk material tensile tests were conducted to investigate the effect of strain rate and silver content on the material properties of Sn–Ag–Cu solders.Solder joint array shear and tensile experiments were conducted on WLCSP speci-mens of different alloy materials under different strain rates and loading orientations to investigate the effects of strain rate,silver content in Sn–Ag–Cu solder joints,and loading orientation on microelectronic packages.Failure analyses were also performed on the fractured dogbone-shaped bulk material test specimens and WLCSP solder joints. 0026-2714/$-see front matter ó2010Elsevier Ltd.All rights reserved.doi:10.1016/j.microrel.2010.01.043 *Corresponding author. E-mail address:mpetanbc@https://www.doczj.com/doc/7717315386.html,.sg (V.B.C.Tan). Microelectronics Reliability 50(2010) 564–576 Contents lists available at ScienceDirect Microelectronics Reliability journal homepage:w w w.e l s e v i e r.c o m /l oc a t e /m i c r o r e l
焊接冶金学(基本原理)课后习题 1.试述熔化焊接、钎焊和粘接在本质上有何区别？ 熔化焊接：使两个被焊材料之间(母材与焊缝)形成共同的晶粒 针焊：只是钎料熔化，而母材不熔化，故在连理处一般不易形成共同的晶粒，只是在钎料与母材之间形成有相互原于渗透的机械结合。 粘接：是靠粘结剂与母材之间的粘合作用，一般来讲没有原子的相互渗透或扩散。 2.怎样才能实现焊接，应有什么外界条件？ 从理论来讲，就是当两个被焊好的固体金属表面接近到相距原子平衡距离时，就可以在接触表面上进行扩散、再结晶等物理化学过程，从而形成金属键，达到焊接的目的。然而，这只是理论上的条件，事实上即使是经过精细加工的表面，在微观上也会存在凹凸不平之处，更何况在一般金属的表面上还常常带有氮化膜、油污和水分等吸附层。这样，就会阻碍金属表面的紧密接触。 为了克服阻碍金属表面紧密接触的各种因素，在焊接工艺上采取以下两种措施：1)对被焊接的材质施加压力目的是破坏接触表面的氧化膜，使结合处增加有效的接触面积，从而达到紧密接触。 2)对被焊材料加热(局部或整体) 对金属来讲，使结合处达到塑性或熔化状态，此时接触面的氧化膜迅速破坏，降低金属变形的阻力，加热也会增加原于的振动能，促进扩散、再结晶、化学反应和结晶过程的进行。 3.焊条的工艺性能包括哪些方面? （详见：焊接冶金学（基本原理）p84） 焊条的工艺性能主要包括：焊接电弧的稳定性、焊缝成形、在各种位置焊接的适应性、飞溅、脱渣性、焊条的熔化速度、药皮发红的程度及焊条发尘量等 4.低氢型焊条为什么对于铁锈、油污、水份很敏感？（详见：焊接冶金学（基本原理）p94） 由于这类焊条的熔渣不具有氧化性，一旦有氢侵入熔池将很难脱出。所以，低氢型焊条对于铁锈、油污、水分很敏感。 5.焊剂的作用有哪些？ 隔离空气、保护焊接区金属使其不受空气的侵害，以及进行冶金处理作用。 6.能实现焊接的能源大致哪几种？它们各自的特点是什么？ 见课本p3 ：热源种类 7.焊接电弧加热区的特点及其热分布？（详见：焊接冶金学（基本原理）p4）热源把热能传给焊件是通过焊件上一定的作用面积进行的。对于电弧焊来讲，这个作用面积称为加热区，如果再进一步分析时，加热区又可分为加热斑点区和活性斑点区； 1)活性斑点区活性斑点区是带电质点(电子和离于)集中轰击的部位，并把电能转为热能； 2)加热斑点区在加热斑点区焊件受热是通过电弧的辐射和周围介质的对流进行的。 8.什么是焊接，其物理本质是什么？ 焊接：被焊工件的材质(同种或异种)，通过加热或加压或二者并用，并且用或不用填充材料，使工件的材质达到原子问的结合而形成永久性连接的工艺过程称为焊接。
SMT焊点可靠性研究 近几年，随着支配电子产品飞速发展的高新型微电子组装技术--表面组装技术（SMT）的 飞速发展，SMT焊点可靠性问题成为普遍关注的焦点问题。 与通孔组装技术THT（Through Hole Technology）相比，SMT在焊点结构特征上存在着很大的差异。THT焊点因为镀通孔内引线和导体铅焊后，填缝铅料为焊点提供了主要的机械强度和可靠性，镀通孔外缘的铅焊圆角形态不是影响焊点可靠性的主要因素，一般只需具有润湿良好的特征就可以被接受。但在表面组装技术中，铅料的填缝尺寸相对较小，铅料的圆角（或称边堡）部分在焊点的电气和机械连接中起主要作用，焊点的可靠性与THT焊点相比要 低得多，铅料圆角的凹凸形态将对焊点的可靠性产生重要影响。 另外，表面组装技术中大尺寸组件（如陶瓷芯片载体）与印制线路板的热膨胀系数相差较 大，当温度升高时，这种热膨胀差必须全部由焊点来吸收。如果温度超过铅料的使用温度范围，则在焊点处会产生很大的应力最终导致产品失效。对于小尺寸组件，虽然因材料的CTE 失配而引起的焊点应力水平较低，但由于SnPb铅料在热循环条件下的粘性行为（蠕变和应力松弛）存在着蠕变损伤失效。因此，焊点可靠性问题尤其是焊点的热循环失效问题是表面组装技术中丞待解决的重大课题。 80年代以来，随着电子产品集成水平的提高，各种形式、各种尺寸的电子封装器件不断推出，使得电子封装产品在设计、生产过程中，面临如何合理地选择焊盘图形、焊点铅料量以及如何保证焊点质量等问题。同时，迅速变化的市场需求要求封装工艺的设计者们能快速对新产品的性能做出判断、对工艺参数的设置做出决策。目前，在表面组装组件的封装和引线设计、焊盘图形设计、焊点铅料量的选择、焊点形态评定等方面尚未能形成合理统一的标准或规则，对工艺参数的选择、焊点性能的评价局限于通过大量的实验估测。因此，迫切需要寻找一条方便有效的分析焊点可靠性的途径，有效地提高表面组装技术的设计、工艺水平。 研究表明，改善焊点形态是提高SMT焊点可靠性的重要途径。90年代以来，关于焊点 形成及焊点可靠性分析理论有大量文献报导。然而，这些研究工作都是专业学者们针对焊点 可靠性分析中的局部问题进行的，尚未形成系统的可靠性分析方法，使其在工程实践中的具体应
金属材料焊接性知识要 点 Document number：NOCG-YUNOO-BUYTT-UU986-1986UT
金属材料焊接性知识要点 1.金属焊接性：指同质材料或异质材料在制造工艺条件下，能够形成完整接头并满足预期使用要求的能力。包括（工艺焊接性和使用焊接性）。 2.工艺焊接性：金属或材料在一定的焊接工艺条件下，能否获得优质致密无缺陷和具有一定使用性能的焊接接头能力。 3.使用焊接性：指焊接接头和整体焊接结构满足各种性能的程度，包括常规的力学性能。 4.影响金属焊接性的因素：1、材料本因素2、设计因素3、工艺因素4、服役环境 5.评定焊接性的原则:（1）评定焊接接头中产生工艺缺陷的倾向，为制定合理的焊接工艺提供依据；（2）评定焊接接头能否满足结构使用性能的要求。 6.实验方法应满足的原则：1可比性2针对性3再现性4经济性 7.常用焊接性试验方法： A:斜Y坡口焊接裂纹试验法:此法主要用于评定碳钢和低合金高强钢焊接热影响区对冷裂纹的敏感性。B:插销试验C:压板对接焊接裂纹试验法D:可调拘束裂纹试验法 一问答：1、“小铁研”实验的目的是什么，适用于什么场合了解其主要实验步骤，分析 影响实验结果稳定性的因素有哪些 答：1、目的是用于评定用于评定碳钢和低合金高强钢焊接热影响区对冷裂纹的敏感性。评定碳钢和低合金高强钢焊接热影响区对冷裂纹的敏感性时，影响结果稳定因素焊接接头拘束度预热温度角变形和未焊透。（一般认为低合金钢“小铁研实验”表面裂纹率小于20%时。用于一般焊接结构是安全的） 2、影响工艺焊接性的主要因素有哪些 答：影响因素：（1）材料因素包括母材本身和使用的焊接材料，如焊条电弧焊的焊条、埋弧焊时的焊丝和焊剂、气体保护焊时的焊丝和保护气体等。 （2）设计因素焊接接头的结构设计会影响应力状态，从而对焊接性产生影响。 （3）工艺因素对于同一种母材，采用不同的焊接方法和工艺措施，所表现出来的焊接性有很大的差异。 （4）服役环境焊接结构的服役环境多种多样，如工作温度高低、工作介质种类、载荷性质等都属于使用条件。 3、举例说明有时工艺焊接性好的金属材料使用焊接性不一定好。 答：金属材料使用焊接性能是指焊接接头或整体焊接结构满足技术条件所规定的各种使用性能主要包括常规的力学性能或特定工作条件下的使用性能，如低温韧性、断裂韧性、高温蠕变强度、持久强度、疲劳性能以及耐蚀性、耐磨性等。而工艺焊接性是指金属或材料在一定的焊接工艺条件下，能否获得优质致密、无缺陷和具有一定使用性能的焊接接头的能力。 比如低碳钢焊接性好，但其强度、硬度却没有高碳钢好。 4、为什么可以用热影响区最高硬度来评价钢铁材料的焊接冷裂纹敏感性焊接工艺条件对热影响区最高硬 度有什么影响 答：因为（1）.冷裂纹主要产生在热影响区；
BGA焊点可靠性研究综述 Review of Reliability of BGA Solder Joints 陈丽丽,李思阳,赵金林（北京航空航天大学,北京100191） Chen Li-li,Li Si-yang,Zhao J in-lin(College of Reliability and System Engineering, Beihang University,Beijing100191） 摘要：随着集成电路封装技术的发展，BGA封装得到了广泛应用，而其焊点可靠性是现代电子封装技术的重要课题。该文介绍了BGA焊点可靠性分析的主要方法，同时对影响焊点可靠性的各因素进行综合分析。并对BGA焊点可靠性发展的前景进行了初步展望。 关键词：有限元；焊点；可靠性；BGA 中图分类号：TN305.94文献标识码：A文章编号：1003-0107(2012)09-0022-06 Abstract:With the development of IC packaging technology,BGA is widely used,the reliability of its sol-der joints has became an important subject of modern electronic packaging technology.In this paper,a common method to analysis the reliability of BGA solder joints is introduced,various parameters which were displayed and the factors of influence on the solder joints,reliability were analyzed simultaneity. Based on above,we have an expectation of development foreground of the reliability of BGA solder joints. Key w ords:finite element;solder joint;reliability;BGA CLC num ber:TN305.94Docum ent code:A Article ID：1003-0107(2012)09-0022-06 0引言 近年来，高功能，高密度，高集成化的BGA封装技术成为主流的封装形式，其焊点可靠性是现代电子封装技术的重要课题。电子封装技术的飞速发展，不断为焊点可靠性的研究提出新课题。传统焊点可靠性研究主要依靠实验，近年来有限元模拟法成为焊点可靠性研究的主要手段；微观显示技术的发展，为分析焊点构成成分变化及裂纹产生，发展提供有力的支持；无铅化进程，针对焊点在不同载荷条件下材料性质成为当前研究的热点；不断涌现出大量新型BGA封装形式，其内部结构，尺寸以及空洞对焊点可靠性的影响有待进一步的研究；板级焊点的可靠性也越来越得到重视。本文主要针对以上几个问题进行综述分析。 1焊点可靠性研究方法 传统的焊点可靠性研究主要依靠实验，随着电子产品的微型化，焊点向着更加微小的方向发展，应用实验方法对其可靠性进行分析面临很大的困难。有限元模拟法[1]，将一个结构分离成若干规则的形状单元，并在空间用边界模型来定义每一个单元就可求解整体结构的位移和应力，利用该方法研究焊点的可靠性也成为热点。 针对单独使用实验方法与有限元模拟方法的局限性，现阶段焊点可靠性的研究多采用实验与有限元模拟方法综合使用的方法。分析方法流程汇总如图1所示。 电子显微技术的发展，使得测试手段多样化发展，检测结果更为准确，对于焊点内部化学成分及结构的变化观察更为直观，能够更好地了解其失效原因，失效部位的形成及发展。下面汇总几种常见的测试方法如表1所示。 2器件级焊点可靠性影响因素 器件封装技术的飞速发展，封装结构，尺寸和材料都发生了较大变化。近年来，专家学者对这类器件级焊点可靠性的影响因素进行了大量研究，下面针对其研究成果进行总结概括。 2.1新型BGA封装结构 2.1.1热增强型BGA 随着电子封装向高密度，薄型化的方向发展，封装的尺寸越来越小，器件的功率越来越大，对芯片的热可靠性提出了更高的要求，为减小热阻，提高热性能，产生了多种热增强型BGA，其主要特点是在BGA封装的底部中间位置(芯片)加有一个散热的铜块或铜片，增加热传导能力，主要用于高功耗器件的封装。其主要结构 作者简介：陈丽丽(1986-)，女，硕士研究生，研究方向为系统安全及可靠性。22