methods has suﬃciently been taken for the Sn–37Pb solderjoints case, where the failure mode was almost always the solder fatigue mode (i.e., cracks at the matrix of the solder material). However, the growth of the intermetallic com- pound in high-temperature dwell time must be considered, because an interface crack may occur when a lead-free solder material is used. In order to investigate the interface fatigue behavior using the isothermal fatigue test method, the specimens were heat-treated before the test to grow the intermetallic compound. In this test, the specimens were aged at 85, 125 and 150 C for 500 and 1000 h and some specimens
A glass-epoxy resin substrate with 0.3 mm diameter Cu pads was prepared. On the surface of each pad, electroless Ni plating was conducted. Ni plating layer thickness was 0.003 mm and the layer contained 8 mass%P. The Cu pads were produced by depositing Au on the Ni layer. Au layer thickness in all cases was 90 or 150 nm at plating times of 10 and 30 min, respectively. For Au plating, a solution was prepared so that Au deposition would take place by sub- stitution and reducing reactions that occur simultaneously. The Ni/Au layer surface was observed with a scanning electron microscope (SEM). The Au layer was removed with removal solution including KCN. This removal solution is used generally for quality control of Ni/Au plating process. And its surface was observed by SEM. A Sn–3 mass%Ag– 0.5 mass%Cu lead-free solder ball 0.35 mm in diameter was prepared as a solder joint by reﬂow soldering in a reﬂow furnace. Peak temperature and holding time over 220 C in the reﬂow process were 250 C and 60 s, respectively. All solderjoints were subsequently exposed to heat treatment at 150 C for 500 h. The ball pull test by cold bump pull (CBP)
Considering the eﬀect of solder volume, solder balls of diﬀerent volume were put onto circular solder pads which are 600 mm in diameter. The standoﬀ height and maximum width of solderjoints after reﬂow, was measured and then compared with simulation results. It can be seen from Fig. 7, that simulation results consist with the experimental measurements very well. Therefore, the accuracy of the program developed in this study is veriﬁed in an accurate range. On the solder pad of the same size, the standoﬀ height and maximum width of solderjoints increase with increasing solder volume. However, simulation results are slightly larger than the experiment measurements, and this may due to the volume shrinkage during the solidiﬁcation of solder.
Abstract– Nowadays, microelectronics plays a key role in our daily lives, such as communication, transportation, embedded systems, medicine etc, so we need to make sure that we can rely on microelectronics systems, while considering the thermal and vibratory fatigue, to make sure of that a lot of methods and simulation were introduced into the process of manufacturing. The reliability and fatigue life prediction of a system is an important problem in product design field, These problems are caused by the fatal flaw of the microelectronic packaging which contain solderjoints, their reliability has a great impact on the reliability of the entire packaging structure. The cause of the fatigue failure of solderjoints is usually caused by the accumulation of thermo-mechanical damages due to the operating conditions of thermal and mechanical shock.
3.4 Microhardness measurement of BGA solderjoints To investigate the changes in mechanical properties of BGA solderjoints, microhardness measurements were performed on the cross-sections of thermal-cycled samples. Nine evenly spaced measurements were made on a 3 3 grid for each sample with a Vickers indentor using a load of 25 g. Figure 9 shows graphically the average microhardness of solderjoints at a corner location of each module (Ring-1 position) in terms of solder composition, cooling rate and ATC scheme. The average microhardness of thermal cycled joints decreases as the ATC condition becomes more severe, such as with increasing T or cycle time in comparison with the as-assembled sample of each group. Except for the solderjoints containing 0.2% Bi, the average hardness gradually decreases as Ag content decreases. It is not obvious to note the eﬀect of cooling rate on the microhardness of solderjoints either as-assembled or thermal-cycled. The microhard- ness measurement was also performed on solderjoints at the center location of each module, where the DNP is zero. The average hardness reduction of the center location was less pronounced than at the Ring-1 position after thermal cycling. This can be understood by recognizing the fact that a solder joint at the center location only experienced thermal exposure, while solderjoints at the corner experienced both thermal exposure and cyclic strain. The latter facilitate microstructural changes as well as crack growth in solderjoints located at the corner.
In this study, the microstructure and shear strength of Cu/solder/Cu solderjoints made from Sn–Ag–Cu and Sn– Ag–Cu–Co alloys were characterized and compared to re- sults for joints made from Sn–3.5Ag to provide a baseline. Shear strength was employed as a microstructure-sensitive mechanical property measurement to provide an indication of joint reliability. Microstructural analysis revealed that the Sn– Ag–Cu near-eutectic alloy exhibits a more highly refined mi- crostructure containing intermetallic particles of both Ag 3 Sn
The Cu/SAC/Cu samples were further subjected to iso- thermal aging at 348 K, 373 K and 423 K in air for 2, 4, 6, 8 and 12 day respectively. Optical metallographic microscope and SEM were used to analyze the interfacial microstruc- tures. The thickness of the IMC layer was calculated by the Image-pro plus software. The average shear strength of sever- al solderjoints (shear velocity was 10 mm/min) was identi- fied by the PTR-1101 bonding strength tester. The shear pro- cess is shown in Fig. 2.
Even though the barrier layer has a comparatively slower reaction rate with the solder, in most of ﬂip-chip processes, reactions still occur at the solder/UBM interfaces. Studies of the interfacial reactions are important, because this kind of information is crucial for evaluating whether the UBM provides enough protection. If the protection is not enough, the solder might react with the IC metal pads and cause the failures of the IC. Since all the Pb-free solders and the conventional Pb-Sn solders are Sn-based low melting- temperature alloys, signiﬁcant atomic diﬀusions take place at room temperatures as well as at the elevated temperatures when the electronic products are in use. Interfacial reactions at the solderjoints occur when the solder and the substrate are both in solid states, in addition, during the soldering process when the solder is molten. Numerous studies can be found regarding the solder/substrate interfacial reactions. 2–6)
Gee et al.  have done the electromigration test for SnAgCu and SnPb solderjoints in a WL-CSP package. The package has 36 solder bumps with 500 µm pitch. The dimension of silicon chip is 1.6 mm×1.6mm×0.5mm. Solder bumps are 0.15 mm in diameter and 0.2 in height. The exterior 20 solder bumps are assumed to connect with each other in a daisy chain as shown in Fig.1. Sub-model technique in ANSYS is introduced to get the better response of the electronic migration. The global thermal-electric coupled field model uses Solid69 element and the global stress model uses Visco107 element for solder bumps and Solid45 element for the remaining parts of the model. Firstly, the global structure is modeled using relative coarse elements firstly. Secondly, a refined thermal-electric coupled field sub-model and a refined stress sub-model with UBM (Al/NI(V)/Cu) layer are then constructed as shown in Fig. 2.
The underﬁlling BGA as an alternative to direct chip attachment for high density packaging technologies have been developed. This paper discusses the thermomechanical and metallurgical eﬀects of underﬁll material and the resulting improvement in board level reliability for underﬁlled BGA assemblies. Finite element analysis (FEA) models were developed to predict the thermal fatigue life of the solderjoints during thermal cycling tests for BGA assemblies without and with underﬁll material. FEA predicted that the stress concentrated in the solder at the crevice between the solder ball and upper substrate was approximately 60 percent of the stress without underﬁll. Subsequently, the predicted fatigue life was as much as 10 times higher for the underﬁlled assemblies. The thermal fatigue failure of BGA solderjoints was also investigated experimentally using thermal cycle testing with subsequent solder joint analysis by scanning electron microscope (SEM) and energy dispersive X-ray (EDX). The experiments revealed that solder joint failure was caused by propagation of cracks that initiated in the solder at the upper interface between the solder ball and copper pad. The fatigue life of the underﬁlled assemblies was about 8 times that of the assemblies without underﬁll. The results showed that the underﬁll material can play an important role in improving board level reliability for BGA solderjoints in harsh environments.
propagation not only of the peripheral crack but also of ﬁne penny-shaped cracks in the bonded region to evaluate the fatigue life of the solderjoints. The numerical simulation of propagation of peripheral and penny-shaped cracks also gives useful information on the estimation of the fatigue life of the solderjoints because the simulation of the peripheral crack predicts the decrease in the apparent bonded region, and also the calculation of the penny-shaped crack predicts the distribution of the penny-shaped crack in the bonded region. In the present work, the shear fatigue testing was carried out with Sn-3.8 mass% Ag-1.2 mass% Cu alloy solderjoints with a thin solder layer of 60 mm in thickness to examine the change in the area of the bonded region in the fatigue process. At required numbers of fatigue cycles, the specimen was taken out from the fatigue tester and subjected to observation through SAM, and the area of the bonded region was estimated from the acoustic image. After fatigue testing, the fracture surface was observed through SEM and the number and the size of ﬁne penny-shaped cracks were measured in the bonded region. The numerical analysis was also conducted to estimate propagation of the peripheral crack and also penny-shaped cracks in the fatigue process.
Abstract— One of major reasons of failure of solderjoints is known as the thermal fatigue. Also, The failure of the solderjoints under the thermal fatigue loading is influenced by varying boundary conditions such as the material of solder joint, the materials of substrates(related the difference in CTE), the height of solder, the Distance of the solder joint from the Neutral Point (DNP), the temperature variation and the dwell time. In this paper, first, the experimental results obtained from thermal fatigue test are compared to the outcomes from theoretical thermal fatigue life equations. Second, the effects of varying boundary conditions on the failure probability of the solder joint are studied by using the probabilistic methods such as the First Order Reliability Method (FORM) and Monte Carlo Simulation (MCS).
for these compounds, and a good part of that which is avail- able was measured with bulk samples whose relevance to the behavior of solderjoints is not at all clear. Our own recent work, including that reported here, was intended to address this problem by conducting high-temperature creep tests on well-characterized solderjoints. In this paper we shall discuss recent results with Sn–3.5Ag, Sn–3Ag–0.5Cu, Sn–0.7Cu and Sn–10In–3.1Ag (mass%), made as thin joints connecting Cu and Ni/Au metallized substrates. As we shall see, the behav-
The microstructure of solder joint under this set of conditions was observed, with the result shown in Fig. 11. The ﬁllet surface showed corrugations, but no cracks were found. The indium phase initially dispersed in the Sn matrix of a ﬁllet and formed an indium oxidized phase near the ﬁllet surface after being left under the above conditions as well as Sn-Zn-based solder. The change in surface conditions is presumed to attributable to the presence of this indium oxidized phase. No generation of the indium oxidized phase was found in the joint interface at that time. Because no indium oxidized phase was found near the ﬁllet surface after the thermal cycle test, the change in the ﬁllet surface in the thermal cycle test and high-temperature/high-humidity test is presumed to occur by diﬀerent mechanisms.
To verify the observed variations in the shear strength, the fracture surfaces after ball shear testing were examined using SEM. Figure 9 shows the cross-sectional and top views of the fracture surfaces of the Sn–9Zn/ENIG joints aged at diﬀerent conditions. The direction of the shear action was from left to right in the cross-sectional views and from top to bottom in the top views. The shear direction is indicated by white arrows. Regardless of the aging conditions, the fractures always occurred in the bulk solder. A similar observation was reported recently in the literature. 10) Date et al. investigated the interfacial reactions and impact reliability of Sn–8Zn–3Bi solder balls (300 mm in diameter) on the Au/Ni–P/Cu layer during aging at 423 K. They reported that all of the Sn–Zn– Bi/Au/Ni–P/Cu joints exhibited bulk fracture at 423 K,
In the case of soldering the Cu block plated with Ni/Au, another soldering process was carried out, as shown in Fig. 2. First, the solder paste was applied on the Cu substrate with a thickness of approximately 60 μm. Then, it was heated using the infrared gold image furnace in a nitrogen atmosphere with a flow rate of approximately 1 L/min at 170C for 1 min to form a thin Sn-57Bi-1Ag solder coating layer. The organic residue of the solder paste was cleaned in an ultrasonic bath with alcohol and a small amount of flux was applied on the thin solder coating. Next, the Cu block with the Ni/Au layer was mounted and a weight of 4.4 g (0.014 MPa) was placed on the Cu block and heated in the same conditions as de- scribed above. The temperature during the heating process was also controlled using the thermocouple attached near the sample, so that the heat loss from the weight was negligible.
From the shear tests, the ultimate shear force required to rupture the solderjoints were recorded and then plotted as a function of field time (in years). For each of the five component types a number of components were sheared for any particular board and the average shear force value was taken. Fig. 4 shows the average shear forces as a function of field time and the deviations for 0603 resistors. The average shear force values were calculated for components in the same way and plotted in fig. 5.The variations in shear force (as in fig. 4) could primarily be due to inconsistency in the amount of solder paste used during reflow soldering. In deed previous research found that stencil printing (used for depositing solder paste) accounts for more than 60% of solder joint assembly defects .
In the SACNG solder, the formation phases in the solder grow up to several-micron meters after heat exposure treatment. Figure 6 shows EPMA mapping analysis results for the cross section of the solder ball joint with the SACNG solder and the Cu pad after heat exposure treatment. Cu, Sn and Ni atoms are detected in dark-gray phases observed in a back-scattered electron image. Since such phases have similar microstructures to that of the reaction layer formed at the joint interface, the phases are probably Cu–Sn compounds with a few atomic percent Ni atoms. In contrast, similar Cu–Sn compounds were scarcely observed in the joint with the SA and SAC solders and the Cu pads. It means that the supplement of a small amount of Ni in the Sn–Ag–Cu solder induces the growth of Cu–Sn compounds in the solder under heat exposure conditions. This is because Ni atoms become the nucleation sites of the formation of Cu–Sn compounds and many nuclei of Cu–Sn compounds are already existed in the solderjoints after reﬂow soldering.
Solderjoints were polished by three kinds of SiC papers (number: 120, 600 and 1200) and by three grades of diamonds pastes (6, 3 and 1 mm). Etching was carried out in hydrochloric vapor at room temperature for several seconds. Then, microstructures of the specimens were observed by an optical-microscope, a scanning electron microscope (SEM) and a transmission electron microscope (TEM). Intermetallic compounds were identiﬁed by electron diﬀraction patterns and by an energy dispersive X-ray analysis (EDX). Crystallographic orientations of solder bumps were identiﬁed by the electron backscattering pattern (EBSP) with an electron beam of about 200 nm in diameter. Measuring step in EBSP measurement was 1 mm. In the present paper, criterion angle of general grain boundaries were over 15 degrees.