Assessing the joints in Surface-Mounted assemblies

Assessing the joints in Surface-Mounted assemblies Dr Nihal Sinnadurai

BPA (Technology and Management) Ltd, Abinger House, Church Street, Dorking, Surrey, UK. Kenneth Cooper

British Telecom Research Laboratories, Martlesham Heath, Ipswich, Suffolk, UK. John Woodhouse

British Telecom Materials and Components Centre, Bordesley Green, Birmingham, UK. With the advent of surface-mounting technology, an apparent problem that has received much attention is the potential failure of solder joints between microcomponents and printed circuit boards. One tequnique used to assess such joints is to subject the assemblies to thermal cycling over a wide temperature range. The validity of the method has, however, not been proven and there are serious reservations about its relevance when considering: the mechanical and thermal properties of solders and other materials associated with the assemblies, the temperature excursions and their rates likely to be encountered in real life, and the more probable problems due to intermittent power dissipation in surface-mounted components. Thus, thermal cycling may prove to be useful only in process development and quality testing. In the event, if the object is to subject the joint to cycling strain, then a more rapid and therefore more efficient test has been devised whereby mechanical cycling is employed to simulate the effects of thermal cycling.

1. Introduction Modern electronics assemblies are increasingly employing surface mounted components (SMCs) soldered on to circuit boards, because of the advantage to be gained in space, cost and electrical performance. However, because many types of SMC are leadless, the solder joints are the sole providers not only of the electrical connection but also the mechanical attachment, and this vulnerability has been exposed by the stresses that can be included in the joints from a variety of possible causes. One such cause, which has been given much prominence, is that arising from the mismatch of thermal coefficients of expansion (TCE) between the SMCs and the substrates (which are usually printed wiring boards (PWBs)). Of course, the advent of any new technology inevitably gives rise to initial problems, and these have to be solved. It is the intention, in this paper, to put the problem of mismatch into perspective and to suggest that testing by thermal cycling is more appropriate as an aid to process development and joint quality assessment than as a test of solder joint reliability. In the event, a simpler mechanical test may be employed. MICROELECTRONICS JOURNAL Vo117 No 2 9 1986 Benn Electronics Publications Ltd, Luton 21

2. Mechanical strains induced between SMCs and substrates

One of the potential causes of mechanical strain in the solder joints of surface-mounted assemblies (SMAs) arises during excursions in temperature, because of differences between the TCEs of the SMCs and substrates, particularly between ceramic chip carriers (TCE 6-7 ppm/ ~ and PWBs (TCE 12-16 ppm/~ Consequently, repeated temperature cycling may give rise to fatigue failure of the joint. The origins of these strains in an idealised solder joint are well known for the case where the substrate expands and contracts at a greater rate then the component. This simple model is the basis for most of the analytical models, which consider only simple shear strains to exist in the joints. However, the solder joints in real SMAs also contain fillets, which add complexity to the strain in the joints through the additional tensile and compressive forces. Hall and colleagues ~have shown, by the use of strain gauges and laser holography, that the actual deformation of solder joints arises from a complex combination of three deformation modes (Fig. 1). They also show that the deformation of the joint during ambient temperature cycling is significantly different from that during power cycling, and conclude that power cycling cannot correctly be simulated by ambient temperature cycling. Similar conclusions were also reached by Engelmaierz. I

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Fig. 2 Schematic fatigue stress or strain cycle 22

3. Fatigue of metals The phenomenon of fatigue is responsible for the majority of mechanical failures in engineering components in service, and is caused by persistent cycling stress of a lower magnitude than the normal monotonic fracture stress of the metal. The low-stress cycling causes cracks which nucleate at the surface and grow into the bulk of the metal until the residual volume is small enough for the applied stress to exceed the monotonic fracture stress, when normal tensile or shear failure will occur. A typical fatigue stress cycle is shown in Fig. 2, and defines the maximum and minimum stresses, the mean stress, and the stress amplitude. It should be noted that the cycle need not be a sine wave, and the mean stress is not necessarily zero- an important factor which is considered later. Because solders are ductile alloys, the fatigue regime of most interest is that which is dominated by high plastic strains, and an empirical relationship, known as the Coffin-Manson relationship 3, has been developed, which attempts to relate fatigue life to the plastic strain amplitude: A E d 2 = Ef.(2Nf).c where/k Er,/2 is the plastic strain amplitude, Ef is the fatigue ductility coefficient, which is the strain intercept at one local reversal (i.e. 2Nf = 1) 2Nf is the number of strain reversals to failure c is the fatigue ductility exponent, a material property normally in the range -0.7 to -0.5.

4. Models for thermal fatigue failure of solder joints 4.1 A basic Coffin-Manson model The Coffin-Manson expression has been used by several workers as a basis for models of the fatigue life of leadless ceramic chip carriers soldered onto PWBs. These models are generally of the same form, and relate the strain to the carrier size, the height of the solder fillet and the temperature cycle range. Such an expression by Taylor and Pedder 4 is: Nf = 2~1

[ 2~/2h ( l + v ) . i [ ~ ](lln+ e ) ] 167

where Nf is the number of cycles tosolder joint failure, h is the height ("stand-off") of each solder pillar, v is Poisson's ratio e is the engineering strain in simple tensile fracture, In ( l + e ) is the natural logarithmic strain, L is the length of an edge of the chip carrier, A T is the range of the temperature cycle, Aa is the difference in TECs between chip carrier and PWB, is an empirical correction factor. This expression is very much a simplificaltion, but shows that the fatigue life of solder joints decreases with increasing size of chip carrier, temperature cycle excursion and TCE difference. The fatigue life can be increased by increasing the stand-off, and by increasing the ductility of the solder alloy. Additional factors not accounted for by the above expression inelude the variation of the properties of solder alloys with temperature and strain rate, and the dependences on the moduli of the chip carrier and substrate materials.

23

4.2 Effects of Solder Joint Geometry on the Strains in the Joints The Coffin-Manson expressions indicate that the geometry of the solder joints have a significant influence on the deformations induced in SM assemblies, and thus will affect the fatigue behaviour of the joints. The variables include the stand-off and the extension of the solder fillet beyond the edge of the component. Increasing the stand-off decreases the shear strain for a given displacement and will increase the fatigue life of the component4. Extending the filldt increases the area of the joint and thus decreases the shear stress in the joint. The effect on the shear force to failure of 44-terminal chip carriers, by varying these parameters, is shown in Fig. 3. The fillet height does not affect these results greatly, but the required force (P) is increased by an increase in extension up to a limiting value of about lmm. The practical effect of extending the fillet from 0 to lmm has been observed 5 to be an increase in fatigue life from 200 to 500 cycles over the temperature range -55~ to +125~ and similar observations are being made in experiments currently in progress. In order to accommodate the linear stresses and the bending moment in the solder joint, finite element analysis has shown6 that the best 50

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shape for the solder joint is with the larger area to be with the chip carrier and a smaller area making well wetted (concave) contact with the PWB, i.e. an inverted fillet. The geometry of the solder joint also influences the heat flow from the component and therefore affects the temperature gradients between a chip carrier and the substrate during power dissipation. Clearly the difference increases with stand-off height and decreases with area of joint. The effect of the stand-off on the substrate temperature for a constant dissipation in a chip carrier is shown in Fig. 4 The above discussion indicates that TCE mismatch can be accommodated by designing a suitable geometry for the solder joints, and the requirement will depend on whether the heating is by ambient temperature excursions or by power dissipation. 4.3 Effects of Strain Rate and Temperature on the Mechanical Properties of Solder Both the monotonic and cycling mechanical properties of solder alloys are dependent on strain rate and temperature. As an example 7, the 0.2% yield stress of 63:37 tin:lead solder at -50~ is five times higher than the value of 125~ In addition, the limit of proportionality (i.e. elastic strain) is known 8 to decrease by a factor of two from -70~ to +60~ The ultimate shear stress of solder joints is also directly proportional to the strain rate and indirectly proportional to temperature, as shown in Fig. 5 for 60:40 solder joints 9"1~ Hence, the fatigue failure in 60:40 solder joints is also directly proportional to the strain rate t~ as shown in Fig. 6. The fatigue strength of 60:40 solder joints decreases with increasing temperature 9"1~ as shown in Fig. 7. In addition, the mean wdue of the stress in a fatigue inducing cycle also has an effect (Fig. 8); longer survival being obtained with zero stress at the minimum of the cycle ~~ As such effects occur within the temperature range commonly employed for thermal cycling tests, distinct differences in plastic strain will therefore be induced at each extreme of a cycle, and will be unique to the particular temperature range and rate of cycling. Such dependences on strain rate and temperature arise because the rate of application of strain is usually so fast that stress relaxation cannot occur in the joints. Stress relaxation itself is related to creep, a high temperature process, in which an applied stress decays with time, converting elastic strain to plastic strain. Most models of thermal fatigue of chip carrier joints ideally assume all the strain to be plastic, whereas the tests derived from the models usually do not allow full plastic strain to develop, and actual practical events depend on strain rate and dwell times at the extremes. The temperature range and mid-point of tile thermal cycle are also crucial, as they determine the maximum and minimum and mean stresses. Fig. 5

Ultimate shear stress of 60:40 Sn:Pb solder joints as a function of temperature and strain rate

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5. Application of Cyclic Tests It is evident from the above arguments, that there are strong reservations about the validity of simulating long slow changes in ambient temperature or even rapidly switched power dissipation in SM assemblies by rapid cycling of the ambint temperature. In addition, the models for joint failure on which such tests are based are themselves inadequate. Because the validity of any accelerated reliability test has to be proven by an ability to extrapolate the test condition to the real operating condition, it follows that thermal cycling is not a valid reliability test of potential solder joint failures. Nevertheless such a test is not without value, because it may be used to reveal dificiencies or improvements in joints by comparison with others, and has been used to advantage in optimising processes". Thus a cycle strain test may be used as an effective test method in process development and in quality control, but not in reliability assessment. As thermal cycling has only a tenuous relevance to real events, the application of cyclic shear strain between SMCs and substrates may be better achieved by more direct means. Therefore, an alternative technique has been devised to replace thermal cycling by direct mechanical cycling~2which is both quicker and easier, and hence is more efficient. 5.1 The Mechanical Cycling Test and its Development According to the planar models discussed in section 4.1, shear strain is induced during thermal cycling of SM assemblies because of the difference in the TCEs between chip carrier and substrate. This shear strain can be equated to linear strains resolved in two orthogonal directions. The alternative mechanical cycling technique may then be used to simulate strain along either orthogonal axis and can be varied over a wide range, both within and beyond the elastic limit. The versatility of the technique allows the test to be used to apply either a typical strain experienced by the assembly or an exaggerated strain for overstress testing purposes. In the development of the mechanical cycling technique, in order to apply linear strain, it was necessary to split the test substrate into two halves and solder the chip carrier to be tested across the split, with opposite rows of its solder joints attached to each half of the substrate. The two edges of the substrate opposite the split were then clamped in the opposing jaws of a general-purpose tensile tester capable of applying tensile, compressive and cyclic strain to the assembly (Fig. 9).

Jaws--~ •L-older Joints

F Fig. 9 CCC/PWB assembly mounted in test jaws

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5.2 Simulation of Thermal Cycling During thermal cycling, the amplitude of the strain experienced by the assembly may be equated to the linear strain induced by linear thermal expansion according to the simple expression: A1 = A a . A T . L where a 1 is the strain amplitude, and the other terms are as defined in section 4. I In order to simulate the strain experiences during thermal cycling, it would be necessary to apply this calculated strain during mechanical cycling. 5.3 Comparison of Mechanical and Thermal Cycling The correlation between mechanically and thermally induced strain has been verified experimentally by conducting both mechanical and thermal cycling tests on similar assemblies. The results of both forms of cycling test yielded the now familiar relationship, plotted in Fig. 10 whereby an increase in strain results in a proportional decrease in the number of cycles to failure. The results of mechanical and thermal cycling are further compared for two different magnitudes of strain (0.018mm and 0.038mm) in Fig. 11. Both sets of results show a similar failure rate, but with different onsets for the mechanical and thermal cycling tests, indicating similar failure mechanisms. Therefore, it appears that mechanical cycling can be used as an alternative, and probably more effective method (because of its simplicity), to thermal cycling for applying cyclic strain to SM assemblies. Strain (turn I

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6. Mechanical Cycling Applications Adjustments to solder joint parameters can lead to the advancement or delay of the onset of their failure. Because of its flexibility and simplicity, the mechanical cycling tests ideal for assessing the effect of such adjustments, It is also useful for assessing the effect of changes in the test parameters, two examples of which are now considered. 6.1 Effect of Cycling Frequency Shown in Fig. 12 is the effect on the beginning of a tensile curve of altering the strain rate during a simple tensile test of an SM assembly. Clearly, as discussed in section 4.3 the greater the strain rate the higher the load and hencethe greater the stress in the solder joints. As the strain rate is directly proportional to the cycling frequency during mechanical cycling, the increased stress induced at a higher cycling frequency would be expected to reduce the number of cycles to failure. The expected result was indeed obtained as shown in Fig. 13. 2B

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Cycles to failure against cumulative percentage failure for mechanical cycling and thermal cycling

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The observations plotted in Fig. 13 were obtained during tests at relatively high frequencies at which the stress was applied without permitting relaxation in the solder joint. Indeed, as discussed in section 4.3, at strain rates at which stress relaxation and creep can occur, the result would be an actual increase in the number of cycles to failure with increasing (but slow) strain rate. Many real-life strain related failures probably do occur within a stress relaxation regime, and would not be correctly simulated either by fast thermal or mechanical cycling (which is even faster). It is therefore paramount that the object of conducting the mechanical (or indeed thermal) cycling tests be ,considered at the outset. 6.2 Effect of Temperature In order to determine the effect of temperature, a resistive heating element was attached to the chip carrier of the assembly to be tested, so enabling the temperature of the assembly to be raised above ambient. Results of tests using this arrangement showed that an increase in temperature of the assembly led to a decrease in the number of cycles to failure, as shown in Fig. 14, confirming the fatigue relationship discussed in section 4.3. 30

6.3 Rapid Assessment The above two examples demonstrate the way mechanical cycling may be used for rapid assessment of changes in test parameters. It is also ideal for assessing changes in assembly construction such as variations in stand-off height, solder pad fillet shape and extensions, solder alloy composition etc. 7. Discussion and Conclusions

It is evident, from this examination of the parameters affecting stress and strain in solder joints of surface-mounted assemblies, that actual failure in joints occurs by a more complex combination of strains than is simulated by thermal cycling which, therefore, is not relevant as an accelerated test for reliability assessment. Nevertheless some form of cyclic test to induce fatigue in solder joints is useful for process development and in quality control. As the rate of thermal cycling is already too fast for stress relaxation to occur, and, as the observations have confirmed that similar fatigue mechanisms are caused by mechanical and thermal cycling, there is advantage to be gained from the more direct mechanical cycling approach to provide rapid testing of non-relaxed solder joints. 8. Acknowledgement

Acknowledgement is made to the Director of Systems Evolution and Standards of British Telecom for permission to publish this paper. 9. References

[i] P.M. Hall, T.D. Dudderar and J.F. Argyle, "Thermal Deformations Observed in Leadless.Ceramic Chip Carriers Surface Mounted to Printed Wiring Boards", Proc IEEE 33rd Electronic Components Conference, pp 350-359, 1983. [2] W. Engelmaier, "Fatigue Life of Leadless Chip Carrier Solder Joints During Power Cycling", IEEE Trans., Components, Hybrids and Manufacturing Technology, CHMT-6, pp 232-237, 1983. [3] S.S. Manson, "Thermal Stress and Low Cycle Fatigue", Mograw-Hill, 1966. [4] J.R. Taylor and D.J. Pedder, "Joint Strength and Thermal Fatigue in Chip Carrier Assembly", ISHM International J. Hybrid Microelectronics, Vol 5., pp 209-214, 1982. [5] R.W. Korb and D.O. Ross, "Direct Attachment of Leadless Chip Carriers to Organic Matrix Printed Wiring Boards", IEEE Trans. Components, Hybrids and Manuf. Technol., CHMT-6, pp 227-231, 1983. [6] W.M. Sherry and J.S. Erich, "Fatigue Life of Fine-Pitch LCCC Solder Joints: Impact of Joint Shape", IEEE VLSI Packaging Workshop, Santa Clara, September 1984. [7] D.E. Riemer and J.D. Russell, "The Optimised Solder Bond for Ceramic Chip Carriers on Ceramic Boards", ISHM Int. J. Hybrid Microelectronics, Vol. 6, pp 217-222, 1983. [8] B.D. Dunn, "The Resistance of Space-Quality Solder Joints to Thermal Fatigue, Part I", Circuit World, Vol. 5, pp 11-17, 1979, and "Part 2", Circuit World, Vol. 6, pp 16-27, 1979. [9] K.R. Stone, R. Duckett, S. Muckett and M. Warwick, "Mechanical Properties of Solders and Soldered Joints", Brazing and Soldering, No. 4 pp 22-27, Spring, 1983. [10] S.E. Muckett, Private Communicati6n. [ll]M.J. Walker, "Surface-Mounted Device, Hybrid Module Technology Trade-offs in Communications Equipment", ISHM Int. J. Hybrid Microelectronics, Vol. 16, pp 27-33, 1983. [12]J. Woodhouse, "A Technique for Assessing the Solder Joint Reliability of SurfaceMounting Components on Substrates of Dissimilar Material", MSc Dissertation, Middlesex Polytechnic, 1983.

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