F. G. Marín, D Whalley, H Kristiansen and Z. L. Zhang, Mechanical Performance of Polymer Cored BGA Interconnects, Proceedings of the 1th Electronics Packaging Technology Conference, 28.
Mechanical Performance of Polymer Cored BGA Interconnects Francisco Guillén Marín 1, David Whalley 1,2, Helge Kristiansen 2 and Zhiliang Zhang 3 1 Wolfson School of Mechanical and Manufacturing Engineering Loughborough University Loughborough, UK 2 Conpart AS Skjetten, Norway 3 Faculty of Engineering Science and Technology Norwegian University of Science and Technology Trondheim, Norway Abstract This paper presents the results from preliminary models comparing the mechanical performance of polymer cored BGA type interconnects with conventional solid solder BGA balls. The Surface Evolver was first used to predict the solder fillet shapes for use in the models. The mechanical behaviour of some candidate polymer spheres were also measured and the results of these measurements used to estimate the properties of the polymer. These results were then used together in the construction of preliminary elasto-plastic finite element models of the interconnects when subjected to cyclic shear loads. The results indicate that the polymer cored balls may potentially provide substantial improvements in solder joint fatigue life, but also indicate the significant effects that the design variables will have on the achievable benefits. Introduction Solder joint failure is widely accepted as one of the most significant reliability hazards for modern electronic systems. Polymer cored BGA balls have been proposed as an alternative to solid solder balls to improve reliability in more demanding application environments [1]. Their potential advantages are dependant on the increased compliance of the polymer core compared with a solid solder ball, thereby reducing the level of stress imposed on the solder joints during both thermal cycling and shock loading due to impacts. The latter is of particular importance for hand held products assembled using lead free solders, which are much more brittle than traditional tin-lead alloys. In addition to protecting the solder joints the increased compliance of the polymer cored interconnects may also reduce the risk of damage to the component and substrate metallisations. It is also anticipated that the increased compliance provided to the interconnect structure by a polymer cored ball may reduce the requirement for under-filling of components in hand held products, and allow adoption of BGA for safety critical applications in harsh environments. A further benefit is more precise control of component standoff height, which may also be greater for a given connection pitch. The polymer cored ball concept is also potentially extendable to chip scale package (CSP) and flip chip (FC) applications. Computational modelling techniques offer cost effective methods to achieving an understanding of the effect of particle and solder joint dimensions and materials properties on the thermal, mechanical and electrical performance and such techniques therefore offer a route to appropriate materials selection for the polymer spheres and their conductive coatings, and for establishing optimum design parameters such as ball diameter, conductive coating thickness, solder pad diameter, and solder volumes. Earlier work by the authors has evaluated the thermal and electrical performance of the polymer cored interconnects [2], whilst this paper concentrates on the mechanical performance. One of the important inputs to such computational models is the materials properties, however the properties of bulk polymers may not be representative of those of small polymeric spheres and this paper will present results of experimental studies of the mechanical performance of some candidate polymer particles. A numerical model has been developed to extract material parameters such as the compression modulus from these measurements, to facilitate successful modelling [3]. Previous experiments have shown that the effective compression modulus of the polymer particle can vary by two orders of magnitude depending on monomer selection and processing parameters [4]. Furthermore the solder fillet geometries depend not only upon the geometric design parameters, but also on the properties of the solder and surfaces wetted by the solder. The Surface Evolver, a computational modelling tool capable of predicting the equilibrium shape of free liquid surfaces, is therefore used to predict the solder fillet shapes for typical polymer cored BGA assemblies. These results are then used together as inputs to preliminary mechanical models of the behaviour of an assembly manufactured using polymer cored BGA balls. A major cause of solder joint failure is thermo-mechanical fatigue, which is induced by cyclic changes in temperature causing differential expansion of the materials within the assembly. These typically impose mainly shear loads on the solder joints and in this study, rather than model the complete assembly of substrate, interconnects and component package, shear loads were applied directly to the interconnects. This greatly simplifies the modelling, whilst still allowing comparison of the polymer cored interconnects with that of conventional solid solder balls. Solder fillet geometry prediction For the work reported in this paper a conventional solder ball has been used as a benchmark. For an assembly using solder balls the weight of the component is generally insufficient to have a significant effect on the solder joint height, and the solder fillet shape following reflow can therefore be approximated as a truncated sphere. For a given 978-1-4244-2118-3/8/$25. 28 IEEE 1 28 1 th Electronics Packaging Technology Conference
volume of solder and diameter of the solder pads it is therefore straightforward to estimate the joint height and diameter. The benchmark geometry was a solder pad diameter of.4mm, an assembled standoff height of.38mm and a sphere diameter of.55mm. For the polymer cored balls the stand off height of the components is determined by the ball diameter, rather than the volume of the solder joints, however the solder fillet shape will be critical to the stress distribution within the assembly. The Surface Evolver [5] is a public domain computer program which may be used to predict the equilibrium shape of liquid surfaces where their geometry is principally determined by surface tension forces and has been used previously in predicting solder joint shapes. Evolver was therefore used here to estimate the solder fillet geometry. The conductor particle size for the polymer cored interconnects determines both the standoff height and the joint diameter and in choosing the particle diameter it must be decided whether to aim for the same sphere diameter as the solder ball BGA, the same standoff height, or some other figure. In this case it was decided to choose the same standoff height of.38mm, as this would be expected to provide a worst case comparison with the solder ball BGA assembly. The other major variables are the solder pad diameter, solder volume and solder wetting angle. The nominal solder pad diameter was.4mm, but was varied between.36mm and.44mm. The solder volume in each solder fillet was chosen to be representative of that of a 1μm thick stencil printed solder deposit over the nominal pad area, but was varied between a 5μm and 175μm thick stencil print. The wetting angles of typical lead free solders have been found to be somewhat higher than for traditional tin-lead based solders. For example Siewert et al. [6] reported the wetting angle of the SnAgCu eutectic to be 32-36 degrees. A nominal wetting angle of 34 degrees was therefore chosen for the model, but this was considered to be a maximum. Figure 1 shows the results from the Evolver models..25.2.15.1.5 (a) Ball surface 5microns paste 75microns paste 1microns paste 125microns paste 15microns paste 175microns paste.5.1.15.2.25.25.2.15.1.5 (b) Ball surface º wetting 1º wetting 2º wetting 34º wetting.5.1.15.2.25 (c) (d) Figure 1. Evolver predictions of solder fillet shapes for (a) nominal geometry and for effects of (b) pad size (c) solder volume and (d) wetting angle.25.2.15.1.5 Ball surface.36mm pad.38mm pad.4mm pad.42mm pad.44mm pad.5.1.15.2.25 Polymer sphere properties measurement Samples of 38mm diameter spheres manufactured using a candidate condensation polymer system capable of withstanding typical soldering temperatures were mechanically tested by compression between two flat anvils at a rate of.5μm/s using an Instron MicroTester. Figure 2 shows the raw test results for a typical sphere. It can clearly be seen that the load versus displacement is non-linear and that at a displacement of about.13mm the load drops substantially, indicating fracture of the sphere. Load (N) 6. 5. 4. 3. 2. 1....5.1.15 Displacement (mm) Figure 2. Mechanical testing results for a candidate sphere The classical theory for indentation of a sphere on a flat surface is based on work by Hertz. However his model assumes that the contact area is very much smaller than the size of the sphere, and that only local deformation is considered. These assumptions are substantially violated during the present experiments, where the deformation of the spheres is more than 3%, and the contact area approaches the size of the sphere. To overcome these shortcomings, Zhang et al. [3], developed a set of explicit equations to describe the stress (σ) versus strain (ε) behaviour of a sphere for large strains. These equations are based on results from a nonlinear geometric analysis using an ABAQUS finite element (FE) model. Strain and stress for the sphere is defined as follows: Δh ε = (1) D 4F σ = (2) 2 πd where Δh is the reduction in height of the particle, D is the undeformed diameter of the particle and F is the applied force. A quarter of the sphere was modelled using the FE mesh shown in figure 3, where the top rigid plate was modelled using a rigid contact surface. A very fine mesh was used in the contact region and a convergence study was performed to ensure that the mesh applied was sufficiently accurate. The results of the model show that for strains in the region of -1%, the following cubic polynomial function can be fitted to the data to determine K, the compression modulus: σ K =.33ε +.99ε2 1.122ε 3 (3) For strains above 1 % the Poisson ratio starts to influence on the results. However, assuming the Poisson ratio to be 2 28 1 th Electronics Packaging Technology Conference
between.2 and.4, the cubic polynomial function below can be used with reasonable accuracy for strains between 1% and 3%: σ K =.667ε +.515ε2 +.5724ε 3 (4) The corresponding E-modulus can then be calculated using the following equation: E = 3K 4 1 ν 2 ( ) (5) where ν is Poisson s ratio, which for these calculations has been assumed to be.38. Figure 4 shows the results of applying this analysis to the data in figure 3 and reveals a progressive softening of the micro-spheres with increasing displacement. The large scatter seen in figure 4 at very small strains, is due to noise in the measurements and failure to accurately determine the point of first contact. For strains larger than a few percent, these effects are no longer significant. Calculated E modulus [MPa] 3 25 2 15 1 5 Figure 3. Finite element mesh used for the modulus analysis.5.1.15.2.25.3.35.4 Strain [-] Figure 4. Calculated Young s modulus for the candidate spheres Fatigue models Low cycle thermal fatigue has long been recognised as a major failure mechanism in the solder joints of surface mounted components. Due to the complexity of the deformation phenomena occurring during a solder joints life, various fatigue models have been proposed. Such models are typically strain based, stress based or energy based [7]. For this study a plastic strain based model using the Coffin Manson relationship has been used to estimate the fatigue life [8]. The Coffin Manson equation used relates the number of cycles to failure, N f, to the plastic strain range, Δε p : m N f Δε p = C (6) where m is fatigue exponent and C is the ductility coefficient. The values of the coefficients used here are those for 99.3Sn.7Cu solder as reported by Pang et al. [9] i.e. m =.973 and C = 21.3. Boundary conditions If it is assumed the component and substrate are rigid whilst the interconnection, such as a BGA ball is very compliant then the relative displacement, δ, between the top and bottom of the interconnection, due to relative thermal expansion, can be easily calculated [1]: δ = ( α PCB αcomponent ) ΔT DNP (7) Where: α PCB = thermal expansion coefficient of PCB α component = thermal expansion coefficient of component ΔT = temperature change DNP = Distance from Neutral Point In reality the elasticity of component and substrate and the stiffness of the interconnects will result in a smaller displacement, but Equation 7 offers a useful first approximation. Equation 7 also indicates that joints that are located on the edge of the component will generally suffer the largest strains, as their DNP is bigger. For array type packages the joint furthest from the neutral point is located near the package corner, i.e. DNP ( L 2) / 2, where L is the length of the package. For a common 2nd level package, such as a BGA, L is around 21mm, and for a first level package, such as a FC, L is up to around 15mm. Worst case thermal displacements for a 1ºC temperature cycle can therefore be estimated as follows: α PCB = 14 1-6 ºC -1 α ceramic = 7 1-6 ºC -1 ΔT = 1 ºC DNP = 14.8 1-3 m δ BGA = 1μm α ceramic = 7 1-6 ºC -1 α silicon = 2 1-6 ºC -1 ΔT = 1 ºC DNP = 1.6 1-3 m = 5μm δ FC These results show that displacements within the range 5 to 1 μm will not be unusual. For the FE models the boundary conditions applied were for the lower solder pad to be fixed and the horizontal displacement calculated above was applied to the other pad, however the upper pad had to be allowed to move freely vertically whilst constrained against rotation, to ensure a pure shear load rather than the simple shear that would occur if the top pad were vertically constrained. 3 28 1 th Electronics Packaging Technology Conference
FEA Model The FEA modelling software used was ABAQUS. To reduce the complexity of the model, 2D meshes were used, which generally used triangular elements and free meshing algorithms. For this study a ball diameter of 38μm with a 5μm thermally induced displacement was used. Only two materials are present in the standard solid solder ball BGA model: the solder ball and two copper pads. On the other hand, the Polymer Core BGA model incorporates six components in total made from three different materials: the two solder fillets, the polymer ball, the copper plating on the ball and two copper pads. The materials properties used for the models are provided in Table 1. The properties of Sn-3.5Ag lead free solder were used for this analysis [6]. As the fatigue life calculations are based on the calculated plastic strain ranges, running the model for a complete displacement cycle is necessary. The model will then provide the strain levels that the joint absorbs during one whole cycle. Figure 5 shows the equivalent plastic strain results for the standard solid solder BGA with solder mask defined (SMD) pads. A non-smd (NSMD) conventional BGA model was also created for comparison. Copper Polymer Solder E (GPa) 15.5 26.2 Poisson s ratio.32.373.36 Yield strength (MPa) 7 22.5 Tensile strength (MPa) 22 26.6 Elongation (%) 45 24 Δε N f, Conventional (SMD).635 34 Conventional (NSMD).387 56 Polymer core BGA.53 413 Table 2. Predicted cyclic plastic strains and resulting fatigue lives Figure 6. High strain area for conventional BGA model Table 1. Materials properties used for FE analysis Figure 7. High strain area for polymer sphere BGA model Figure 5. Conventional BGA model equivalent plastic strain results Figure 6 shows the critical plastic strain area for the conventional BGA, while figure 7 shows the critical area for the polymer core BGA. Both models suggest crack propagation will start in the solder. However, in the polymer core model the mostly highly strained region is displaced from the interface between the pad and the solder to the solder to ball interface. The predicted cyclic plastic strains from the FEA models and the resulting fatigue life predictions are presented in table 2. Solder fillet geometry sensitivity analysis The shape of the solder fillet is expected to have a significant effect on the polymer cored BGA interconnection reliability and an understanding of the effects of design variables will aid in optimization of the polymer cored interconnections. The co-ordinates of points along the solder fillet curves as previously calculated using Evolver were extracted and used within ABAQUS to generate a series of splines defining the different solder fillet shapes. For these models a simpler elastic/perfectly plastic material behaviour was simulated, rather than incorporating a hardening law, and a higher modulus value for the polymer, representative of bulk materials, was used. These results are therefore not directly comparable with those in the previous section. 4 28 1 th Electronics Packaging Technology Conference
Effect of solder pad size The solder pad size was varied between a diameter of 36μm and 44μm, whilst maintaining the solder volume constant. A roughly linear relationship between pad size and cyclic plastic strain was observed, as shown in figure 8. Strain(%),2 17 18 19 2 21 22 Pad radius(μm) Figure 8. Cyclic plastic strain versus pad radius Effect of solder volume Five different solder volumes were modelled whilst retaining the same nominal pad size of 4μm diameter. As mentioned earlier these solder volumes were representative of those for printing onto the pad area through stencils between 5μm and 175μm thick. If a 2μm thick stencil were to be used the solder volume would be sufficient for the two solder fillets to touch each other around the centre line of the joint, thus merging and creating a single column of solder surrounding the ball. According to the results, as shown in figure 9, the optimum volume seems to be achieved at around a 125μm stencil thickness, whilst for greater solder volumes the strain starts to rapidly increase towards the levels for standard BGA joints. Strain(%),6,55,5,45,4,35,3,25,16,14,12,1,8,6,4,2 3 8 13 18 Volume (stencil thickness(μm)) Figure 9. Effect of solder volume on cyclic plastic strain Effect of wetting angle Previous studies of solder joint reliability have shown that smaller wetting angles will generally be better, by reducing localised stress concentrations at he solder to metallisation interfaces. The wetting angle achievable in a particular situation depends upon many factors such as surface composition and roughness, solder alloy type, flux type and soldering process conditions, and very low wetting angles are not always attainable or predictable [1]. However, according to these FE modelling results, the wetting angle has no major effect on the plastic strain range for the polymer BGA assemblies, as is shown by the cyclic plastic strain results in Table 3. These show that wetting angle is not a critical parameter, although the results do suggest that, given the severe strains occurring at the solder fillet edge, a very small wetting angle is not ideal in this situation as it causes a greater solder fillet curvature and therefore results in a higher stress concentration. Wetting angle Δε º.41 1º.33 2º.41 Table 3. Effect of wetting angle on plastic strain range Discussion of results Figure 1 summarises the predicted solder joint fatigue lives for all of the BGA configurations discussed in this paper, which have been calculated using Equation 6 from the cyclic plastic strains predicted by the FE models. The Coffin- Manson coefficients corresponding to the 99.3Sn.7Cu solder alloy were used in this study. On the other hand, the solder mechanical properties used for the FE models were those for Sn-3.5Ag. Although this may not be consistent, reliable data for lead free alloys are still difficult to find and, as the same parameters were applied to both conventional and polymer cored interconnects, it is considered reasonable to use them for comparison. It is clear from figure 1 that the polymer cored BGA balls potentially offer a substantial solder joint life improvement over the conventional BGA balls, whether SMD or NSMD, that may exceed a factor of ten. Figure 1 also demonstrates that a wide variation in life may occur depending upon the detailed design of the interconnect, with credible variations in solder volume alone resulting in a factor of five lifetime variation. Other significant parameters not addressed in this paper include the sensitivity to the polymer Young s modulus and the type and thickness of the metallic coating on the polymer. A further potentially important factor so far not addressed is the effect of variations in geometry between the top and bottom solder fillets. One other important factor to consider is potential fatigue failure of the copper metallisation on the polymer spheres. In some of the models stresses in excess of the yield point of the copper have been observed. These stresses are expected to increase for more flexible polymers and a trade-off between solder joint life and copper fatigue life is likely to be required in order to optimise overall life. 5 28 1 th Electronics Packaging Technology Conference
Polymer core BGA Conventional (NSMD) Conventional (SMD) 175μm thick stencil 15μm thick stencil 125μm thick stencil 1μm thick stencil 75μm thick stencil 5μm thick stencil.22mm pad radius.21mm pad radius.19mm pad radius.18mm pad radius 2º wetting angle 1º wetting angle º wetting angle 1 2 3 4 5 6 7 8 9 1 Number of cycles to failure Fig 1. Comparisons of predicted life for different BGA configurations Conclusions The results of this work indicate a promising future for polymer cored BGA type interconnections in terms of their mechanical performance. However the results indicate that the achievable performance will be highly sensitive to a number of design parameters and a better understanding of the effects of these parameters will therefore be required in order to ensure the optimum design for specific applications. Also, the results depend strongly upon material properties which are difficult to establish for the polymer particles and further work on measurement techniques is necessary, including extension to measurement throughout the anticipated range of application temperatures for the materials. Future work will explore these issues in more detail and will also evaluate the shock/impact performance of the polymer cored interconnects in comparison with conventional interconnects. 4. Kristiansen, H., Redford, K., Zhang, Z. He, J. Y., Fleissner, M. and Dahl, P. I. Development and Characterisation of Micrometer Sized Polymer Particles with Extremely Narrow Size Distribution, 12 th IEEE International Symposium on Advanced Packaging Materials, ISSN 189-819, 3-5 October 27, San Jose, pp 13-134 5. Brakke, K. The Surface Evolver, Experimental Mathematics, Vol. 1, No. 2 (1992), pp 141-165 6. Siewert, T., Liu, S., Smith, D.R. and Madeni, J.C. Database for Solder Properties with Emphasis on New Lead-free Solders National Institute of Standards and Technology, Colorado, 23, Release 4. 7. Zhang, L., Sitaraman, R., Patwardhan, V., Nguyen, L. and Kelkar, N. Solder joint reliability model vath modified Darveaux's equations for the micro smd wafer level-chip scale package family Proceedings of the 53rd IEEE Electronic Components and Technology Conference, New Orleans, May 27-3, 23, pp572-577 8. Lee, W.W., Nguyen, L.T. and Selvaduray, G.S. Solder joint fatigue models: review and applicability to chip scale packages. Microelectronics Reliability Volume 4, Issue 2, 28 February 2, pp231-244 9. Pang, J.H.L., Xiong, B.S. and Low, T.H. Low cycle fatigue study of lead free 99.3Sn.7Cu solder alloy, International Journal of Fatigue, Volume 26, Issue 8, August 24, pp865-872 1. Li, Y. and Mahajan, R.L. CBGA Solder Fillet Shape Prediction and Design Optimization Journal of Electronic Packaging, Volume 12, Issue 2, June 1998, pp. 118-122 Acknowledgments The authors gratefully acknowledge the support given by Pradeep Hegde from Loughborough University to the particle compression experiments. References 1. Movva, S. and Aguirre, G. High Reliability Second Level Interconnections Using Polymer Core BGAs, Proceedings of the 54th IEEE Electronic Components and Technology Conference, 1-4 June, 24, Volume 2, pp1443 1448 2. Whalley, D.C., Kristiansen, H. and Guillen Marin, F. Thermal and Electrical Modelling of Polymer Cored BGA Interconnects, Proceedings of the 2nd IEEE Electronic Systemintegration Technology Conference, Greenwich, 1-4 Sept, 28, Volume 2, pp 19-116 3. Zhang, Z.L., Kristiansen, H. and Liu, J. A method for determining elastic properties of micron-sized polymer particles by using flat punch test, Computational Materials Science, Volume 39, Issue 2, April 27, pp 35-314 6 28 1 th Electronics Packaging Technology Conference