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Solder-joint reliability of HVQFN-packages subjected to thermal cycling
Microelectronics Reliability 49 (2009) 331–339
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Microelectronics Reliability journal homepage: www.elsevier.com/locate/microrel
Solder-joint reliability of HVQFN-packages subjected to thermal cycling J. de Vries a,*, M. Jansen a, W. van Driel b,c a b c
Philips Applied Technologies, High Tech Campus 7, 5656AE Eindhoven, The Netherlands IMO Back End Innovation, NXP Semiconductors, Gerstweg 2, 6534AE Nijmegen, The Netherlands Delft University of Technology, Mekelweg 2, 2628CD Delft, The Netherlands
a r t i c l e
i n f o
Article history: Received 3 October 2008 Received in revised form 11 December 2008 Available online 4 February 2009
a b s t r a c t In this work experimental results of thermal cycling tests on HVQFN-packages mounted on printed circuit boards are combined with finite element analyses. Validating the finite element analyses by a selected series of small, medium and large HVQFN-packages assembled on printed circuit boards, allows us to determine the performance of this family. To be able to do that, the discriminating parameters that determine the board level performance of this family need to be understood. The emphasis is on the fatigue life of the soldered interconnections as it is influenced by the thermal stress load, the board thickness, and the dimension of the package. Data from different experimental set-ups are compared. An important parameter in this respect is the inclusion of the base material of the panels. The test loads were set to cycling at 40 °C/+125 °C and 20 °C/+100 °C. The results prove that the essential physical properties governing the fatigue life are the stiffness of the complete assembly and the thermal expansion mismatch between the parts. Ó 2008 Elsevier Ltd. All rights reserved.
1. Introduction Over the past few years new types of IC-packages have been developed, among which the quad flat non-leaded (QFN) or micro-lead frame (MLF) packages have become very popular. In this class a variety with a low building height and a heat sink is usually called HVQFN. The advantages of these packages are manifold. They use less board space than other similar packages since they are leadless. Having no solder bumps and leads the electrical properties, such as the self-inductance, are better. The comparatively large heat sink makes them thermally superior. Still, there are some concerns as to the board level reliability. The low solder joint height of typically a few tens of micrometers leads to a much larger shear strain caused by thermal expansion mismatches. And thus the solder fatigue life is reduced. Already some, yet not very much, data have been published on the board level endurance of QFN-type of packages. In one of the earliest papers both the mechanical and the thermal solder joint reliability of bump chip carrier- and QFN-packages was compared [1]. An extensive overview was published that dealt with design parameters of the QFN-package and of the printed circuit board [2]. These comprise such items as the dimension of the solder lands, the height of the interconnection, die size, board thickness, and of course the test condition. One of the first numerical simulation studies that was carried through
* Corresponding author. Tel.: +31 40 27 48765. E-mail address: [email protected] (J. de Vries). 0026-2714/$ - see front matter Ó 2008 Elsevier Ltd. All rights reserved. doi:10.1016/j.microrel.2008.12.007
basically covered the same topics as above [3]. Somewhat similar work combining experiments with modeling was issued in the last few years [4–6]. Yet in all these valuable studies the basic material properties of the printed circuit board were more or less taken for granted. To date one has just begun to address this point [7–9]. After measuring the board material properties the durability of several components – but no QFN – under vibrational or thermal cycling load was simulated. It is the aim of the present work to show the importance to know the board properties. To this end results from various experiments have been evaluated and completed with additional tests. Numerical simulations serve to support the analyses. Whenever possible, analytical models are formulated to illustrate the problem. 2. Experimental issues and simulation For the present work small-sized HVQFN24, medium-sized HVQFN48, and large-sized HVQFN72-type packages were selected as test carrier. In Fig. 1 the layout of such package is shown. They were made with daisy chains to facilitate on-line monitoring of the integrity of the soldered interconnections. In Table 1 the relevant data of the test packages are listed. For all experiments printed circuit boards with four Cu-layers (35 lm) and FR4 were used, the only difference being their layout or thickness. The latter was achieved by adjusting the thickness of the FR4-layer. The panels had NiAu-finish. Each panel could host 30 products.
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In order to capture the actual moment of failure, on-line monitoring of the daisy chain resistance is required. Event detectors serve to detect transient interruptions in the chain. In the present case AnaTech 128/256 STD equipment was used. A failure was defined according to IPC9701 as the first interruption of the daisy chain followed by at least nine repeated events [10]. The threshold value of the daisy chain resistance was set to 1000 X. Commercially available statistical software (Reliasoft, Weibull++ version 6) served to analyze the failure distributions. Weibull or log-normal statistics are commonly used to evaluate these. In the present work Weibull analyses were carried out. The failure distributions are all given with 95% confidence levels; the results are summarized in the lower part of Table 2. In some
Fig. 1. Schematic layout of an HVQFN (left) and back side view on actual sample (right).
The packages were dried for 24 h at 125 °C prior to reflow soldering. Reflow was done at a peak temperature of 250 °C as was determined on the boards. Standard lead-free solder paste (Sn3Ag0.5Cu: SAC305) was applied. For the investigation of the effect of the stand-off, additional solder paste was stencil printed on HVQFN48-packages. The products were reflowed and then the normal assembly process was run. After assembly, the boards were inspected by means of X-ray and cross-sectioning. In Fig. 2 one finds selected pictures of the X-ray inspection of packages immediately after assembling. With some effort even the daisy chains can be discerned. We mention here that the voids in the diepad are typical for the various assemblies studied in this work. In the further analyses voiding was not taken into account. Cross-sections of such packages are shown in Fig. 3. Then the assemblies were subjected to a thermal cycling test in a one chamber system (ESPEC ENX12-7.5 CWL). The test condition was 40 °C/+125 °C with a cycle time of 1 h. In Fig. 4 typical temperature profiles as recorded on the packages are shown, including the profile that was used for the thermal testing at 20 °C/+100 °C/ 1 h-cycle (one chamber Grenco GTTS 125.20S). The temperatures were measured by thermocouples placed inside the soldered joints. Various board level test standards advice to test at least 30 but preferably more products [10]. Table 2 contains the basic elements of the experimental set-up.
Table 1 Package geometry HVQFN. The row ‘‘Experiment ID” refers to Table 2. HVQFN Size (mm) Thickness (mm) Die pad (mm) Die (mm) Coverage (%) Pitch (mm) Experiment ID
24 44 0.85 2.8 2.8 2.0 2.0 51 0.5 B
48 77 0.85 5.3 5.3 2.2 2.2 17 0.5 A, B
48 77 0.85 5.3 5.3 3.5 3.5 44 0.5 C
72 10 10 0.85 6.2 6.2 4.4 4.4 51 0.5 B
Fig. 2. X-ray of assembled packages. Pictures are NOT to scale. (a) HVQFN24. (b) HVQFN48. (c) HVQFN72.
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cases early failures were found, which were suspended from the analyses. Together with the surviving samples these data points are marked with separate symbols in the failure distributions. Numerical simulations were carried out on a quarter model of an assembled package. The moulding compound was simulated with a viscoelastic material model, the solder with a creep model. The finite element suite MARC was used for the modeling. Further details are given in the appropriate sections.
16 14
E' (GPa)
12 10 8 6 4 2 0 0
50
100
150
200
T (°C) Fig. 5. Storage modulus of pcb’s (DMA). See Table 2: A (), B (), C 1.6 mm (—), C 0.8 mm ( ). Fig. 3. Cross–sections of assembled packages. Pictures are to scale; the white scale bar is 3 mm. (a) HVQFN24. (b) HVQFN48. (c) HVQFN72. Half-etched die pad is visible.
length change (ppm)
2000
1500
1000
500
0 0
50
100
150
T (°C) Fig. 4. Temperature profiles 40 °C/+125 °C (solid) and 20 °C/+100 °C (dotted) as measured on the packages.
Fig. 6. Relative length change of pcb’s (TMA). See Table 2: A (), B (), C 1.6 mm (—).
Table 2 Assembly and test parameters: profile, label (ID). Pcb: thickness (t), elastic modulus (E), glass transition temperature (Tg), expansion coefficient (a). Solder joint height (h). Number of samples (N). Weibull-parameters: scale (g) and shape (b); one- (1-p) and two-parameter (2-p) analysis. Open cells: data not determined. HVQFN
24
48
Experiment ID
B1
A
B2
C1
C2
C3
C4
B3
T-profile (°C) t (mm) E (GPa) Tg (°C) a (ppm/K) h (lm) N Analysis type
40/125 1.6 11 174 25.0
40/125 1.6 11 132 13.8 54 60 2-p 3110 4.17
40/125 1.6 11 174 25.0
40/125 1.6 15 131 23.8 28 30 2-p 810 4.17
40/125 0.8 15 125
20/100 1.6 15 131 23.8
20/100 0.8 15 125
40/125 1.6 11 174 25.0
30 2-p 1840 2.98
30 2-p 3090 4.26
30 2-p 7270 6.42
40 2-p 4670 4.01
g b
40 1-p 8400 5.0
72
40 2-p 5520 5.32
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Fig. 7. Set B2: left 100 lm-bar; right 60 lm-bar.
3. Results The elastic modulus and thermal expansion coefficient of complete board stacks including the copper layers were determined. Pieces of the various printed circuit boards were cut out to determine these properties by means of Dynamical Mechanical Analysis (DMA) and Thermal Mechanical Analysis (TMA), respectively. The results of measuring the thermo-mechanical board properties are collected in Figs. 5 and 6 (see also Table 2). Although only FR4-based panels were used in the experiments, there is an appreciable variation in the mechanical properties of the boards. They are all four-layer panels with 35 lm thick Cu-layers. In the boards made from the same material but of different thickness (types C in Table 2) the elastic modulus is about 15 GPa below the glass transition temperature (Tg) which lies around 130 °C. The stiffness of the 1.6 mm pcb is of course higher than for the corresponding 0.8 mm panel. Fig. 5 also shows two other sets of data (A and B): these particular pcb’s are of the same thickness, and both have a lower elastic modulus of about 11 GPa below Tg. However, their thermal expansion coefficients differ considerably, as do their respective glass transition points. All of these differences will undoubtedly have effect on the fatigue life of the assemblies. A higher elastic constant or a larger expansion mismatch must induce more stress during thermal cycling. At several moments during the tests samples were taken out for failure analysis. The failure mode was assessed from cross-sections made of failed samples. In Figs. 7 and 8 some examples clearly reveal solder fatigue. Plastic deformation of the soldered joints and cracks within the solder appear. 4. Discussion of experimental results Per each variation the results will be discussed and compared with literature data whenever these are available. In a few cases
we will use analytical models to illustrate the problem and the advantages of the numerical approach. The technical details of the finite element simulations will be presented in Section 5. The basic elements of the analytical model are given here. For comparing the effects of thermal cycling ranges on the fatigue lifetime (s) one applies the Coffin–Manson relation:
2s ¼ ðDc=2ef Þ1=c ;
ð1Þ
where Dc is the maximum shear strain between the component and the board, ef the fatigue ductility coefficient, and c the ductility exponent respectively. For eutectic SnPb-solder the exponent c is about minus one half, leading to the famous quadratic dependence of the fatigue life on the temperature range. A useful equation to estimate the ductility exponent was given by Engelmaier [11]:
c ¼ c0 þ c1 T center þ c2 lnð1 þ 360=tdw Þ:
ð2Þ
The coefficients ci are specific to the solder material. Further Tcenter is the center temperature of the thermal cycle range, and tdw is the hold time at the extreme temperatures. For three solder alloys the values of ci and the ductility coefficient ef are listed in Table 3 [12,13]. The following equation describes the maximum shear strain (Dc):
Dc ¼ ðDL=hÞDaDT:
ð3Þ
Table 3 Coefficients ci for ductility exponent and ductility coefficient ef for three solder alloys [12,13].
c0 c1 (1/K) c2
ef
SnPb
SAC305
SnAg
0.502 7.34E04 1.45E02 2.25
0.347 1.74E03 7.83E03 3.47
0.416 2.10E-03 1.40E-02 2.25
Fig. 8. Left: set A 100 lm-bar; right: C1 60 lm-bar.
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99.
cumulative failures (%)
90.
50.
10. 5.
1. 100.
1000.
10000.
N-cycles Fig. 9. Weibull distributions. Temperature profile: ( , C1) 40 °C/+125 °C, ( , C3) 20 °C/+100 °C. Pcb 1.6 mm. Labels see Table 2.
L is the span of the soldered joints under consideration, DL their thermal displacement, h is their height, Da is the thermal expansion mismatch between the component and the board, and DT is the temperature difference to which the assembly is subjected. For the fatigue life as function of the solder joint height one has:
s h1=c
ð4Þ
In case of SAC-solder the proposed data in Table 3 lead to a value for the ductility exponent of c = 0.4. 4.1. Variation of test condition
4.2. Effect of solder joint height The influence of the height of the soldered interconnections on the fatigue life is given in Fig. 10. All data were obtained on boards of 1.6 mm thickness (see Table 2, sets A and C1). The smaller standoff of 28 lm leads to a mean time to failure of 810 cycles. By doubling the solder joint height the fatigue lifetime increases to 3110 cycles, which is a factor of 3.8. Other experimental results for comparison could not be found. But one generally acknowledges the beneficial effect of a larger stand-off.
99.
99.
90.
90.
cumulative failures (%)
cumulative failures (%)
In Fig. 9 the Weibull failure distributions of assemblies subjected to two different test conditions are shown. A summary of
the statistical evaluation is listed in the lower half of Table 2 (test runs C1 and C3 with 1.6 mm thick panels). In the first place it must be noted that both distributions run in parallel, as also the shape parameters (b) show. In the second place from the failure analysis – not shown here – the same failure mode was found. This is a strong indication that both tests invoke the same failure mechanism. The milder temperature cycling condition of 20 °C/ +100 °C leads to almost a fourfold longer fatigue life (3090 cycles compared to 810 cycles). In the series with the 0.8 mm thick boards this is the same: 7270 versus 1840 cycles. Astonishingly little reference data for QFN-packages have appeared in the literature dealing with the temperature cycle range. Syed and Kang [2] mention experiments on QFN32-packages of 5 5 mm2 subjected to three different thermal cycling tests. However, only the moment of first failure is reported as the experiments were still running. Cycling at 40 °C/+125 °C/1-h the first failure is observed at 3352 cycles, at 55 °C/+125 °C/40-min this is at 2610 cycles, while after 5813 cycles testing at 0/+100 °C/30min no failures had occurred yet. The same type of data are listed in an application note of Freescale [14]. For QFN48-products of 7 7 mm2 body size cycling at 40 °C/+125 °C/1-h leads to the first failure at 1340 cycles, and at 0/+100 °C/30-min this increases to 3950 cycles. The temperature ranges in our experiments have a ratio of 1.4, which results in an acceleration factor of 2.3. This must be compared to a ratio of about 4 between the experimental lifetimes. Clearly the relatively simple Coffin–Manson approach does not suffice in the current situation.
50.
10. 5.
1. 100.
1000.
10000.
N-cycles Fig. 10. Weibull distributions. Solder joint thickness: ( , C1) 28 lm, ( , A) 55 lm. Pcb 1.6 mm, 40 °C/+125 °C. Labels see Table 2.
50.
10. 5.
1. 100.
1000.
10000.
N-cycles Fig. 11. Weibull distributions. Pcb thickness: ( , C1) 1.6 mm, ( , C2) 0.8 mm. 40 °C/+125 °C. Labels see Table 2.
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99.
was 1254 cycles. In the work of Mercado et al. [4] the relative effect of the panel thickness on the solder fatigue life of BCC32-packages is given. Going from 0.51 to 1.02 mm thickness the lifetime decreases by a factor of 2.8, and from 1.02 to 1.57 mm by a factor of 1.6. In the present case, the assemblies on the thinner printed circuit boards have a fatigue life that is a factor 2.3 longer than the series on the thick boards. Whereas this agrees fairly well with the referenced literature data, an analytical model is much more difficult to give. Straightforward mechanics tells us that the stiffness of a thin slab is proportional to the cube of its thickness. Considering the assembly as a spring system, and assuming that the energy contained in this system is inversely proportional to its lifetime, only a very qualitative description of the effect of board thickness on fatigue life can be obtained.
cumulative failures (%)
90.
50.
10. 5.
4.4. Effect of package size
1. 100.
1000.
10000.
N-cycles Fig. 12. Weibull-distributions. Package size: ( , B3) HVQFN72, ( , B2) HVQFN48, ( , B1) HVQFN24. Pcb 1.6 mm, 40 °C/+125 °C. Labels see Table 2.
From Eq. (4) it follows that the stand-off change from 28 lm to 55 lm should lead to a 5.7-fold increase in fatigue life. However, not only the solder joint height, but also the thermal expansion coefficient of the printed circuit boards used in the two experiments were different. Taking this fact into account the analytical result becomes even worse. 4.3. Effect of stiffness of printed circuit board The stiffness of the printed circuit boards depends on the material constant – the elastic modulus – and the construction – the thickness – of the board. Experimental results for boards of thicknesses ranging from 0.8 to 1.6 mm are shown in Fig. 11 (data sets C1 and C2 of Table 2). Similar experimental results were reported by Syed et al. [2] on QFN72-packages of 10 10 mm2 size subjected to 40 °C/+125 °C cycling. On boards of 1.6 mm thickness the fatigue life was 854 cycles, whereas for 0.8 mm boards this
To investigate the effect of the package body size on the thermal fatigue life, three different HVQFN-types were chosen (see Table 2, series B). However, after 5400 cycles only the two larger packages had a failure rate that allowed for a statistical analysis, the smallest had only four failures. In order to include at least a good estimate for HVQFN24, the width of the distributions (b, Table 2) for the two larger packages was taken as the starting point for a one-parameter Weibull-analysis of the small package: b = 5. In Fig. 12 we show the graphical result. Thus the scale parameters of the three HQVFN-body sizes can be compared. For HVQFN24 the uncertainty is in the range of ±2000 cycles. As for HVQFN72 it must be noted that the lead frame is half-etched which will increase the compliance of the package and hence the lifetime is expected to be longer. Tee et al. [3] report on board level solder joint reliability of QFNpackages with a different outline. These were tested by cycling at 40 °C/+125 °C with a cycle period of 40 min. The larger QFN52 of 8 8 mm2 has a fatigue life in the range of 631–1426 cycles. For the smaller QFN20 of 4 4 mm2 this is 2743–4894 cycles. The ranges depend on design variations. 5. Numerical model The discussions of the experimental results prove the shortcomings of analytical calculations to analyze the results. In this section
Fig. 13. Quarter FE-model (left), die pad (top right), cross-section showing also die and lead frame (bottom right).
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150
1.E-03
8.E-04
6.E-04 50
T (°C)
strain (-)
100
4.E-04
0 2.E-04
-50
0.E+00 0
2000
4000
6000
8000
time (s) ), and plastic strain (left axis,
we will briefly discuss the finite element model and the results of using it to a parameter sensitivity analysis. The creep model proposed by Darveaux was used to simulate the behavior of the solder material [15]:
n dccr GðTÞ s eQ =kT : sinh h ¼C T GðTÞ dt
ð5Þ
Here dc/dt is the creep strain rate, G is the shear modulus, h the stress level where the power law breaks down, s the applied stress, n the stress exponent, and Q is the activation energy. For Sn3.5Ag the following values were taken: C = 0.4539, G(T) = 19,300 69(T 273) [MPa], h = 1500, n = 5.5, and Q = 0.5 eV [15], see also [16,17]. The finite element model consists of a quarter of the complete assembly (see Fig. 13). Two thermal cycles have been simulated (see Fig. 14). The strain was determined by averaging over one complete soldered connection at the corner of the package since this is the most critical joint. In Fig. 14 the plastic strain and the creep strain are shown as they build up during thermal cycling. Plastic strain is only 10% of the creep strain. As the temperature is at its lower extreme, the creep mechanism is hardly active, while at the higher extreme only little creep occurs because the stress is much lower. Only during the temperature ramps an appreciable amount of creep strain develops. Application of the finite element model to various assemblies leads to a satisfactory result as depicted in Fig. 15. Here, one finds the lifetime simulated for the B-and C-series and calibrated against the respective experimental results of HVQFN48 at 40 °C/+125 °C on 1.6 mm boards (see Table 2, series C1 and B2). One recognizes the outcome of testing at a different cycling condition which was discussed in Section 4.1 (Table 2, series C3 and Fig. 9). Likewise, the tests carried out on packages soldered onto a thinner board, discussed in Section 4.3, are shown (Table 2, series C2 and Fig. 11). To further validate the model, also other HVQFN-packages were modeled and compared to experimental test results. These are the smaller HVQFN24 (4 4 mm2) and the larger HVQFN72 (10 10 mm2) which can be found in Table 2 series B1, B3 and Fig. 12. Finally, the test series with a larger solder height (Table 2, series A) was simulated. As one can see in Fig. 15 this result does not
). (For interpretation of the references in color in this figure legend,
compare very well to the experiment. The main difference is the much lower thermal expansion coefficient of the printed circuit board used for this experiment (see Fig. 6 and Table 2). Still, keeping this in mind, one may safely conclude that the numerical model is sufficiently reliable to proceed with the sensitivity analysis. The analyses of the experiments sofar lead to the notion that the thickness, the elastic modulus, and the thermal expansion coefficient of the printed circuit boards govern the fatigue life of the interconnections and variations therein should not be neglected. In this case one thermal cycle was modeled. Three board parameters were varied (in parentheses the nominal value): – Thickness 0.5–2.5 mm (1.6 mm). – Elastic modulus 10–20 GPa (17.5 GPa). – Thermal expansion coefficient 10–25 ppm/K (17.6 ppm/K). Again, one of the corner joints was used to average the strain that was translated to fatigue lifetime by means of the Coffin–Man8000
simulation (cycles)
Fig. 14. Temperature profile (right axis, red line), creep strain (left axis, the reader is referred to the web version of this article.)
B2
6000
B1
4000
B3 C3
2000
C1 C2
A
0 0
2000
4000
6000
8000
10000
experiment (cycles) Fig. 15. Simulated lifetime (63%-failures) against experimental data (with 95% confidence levels). Data labels refer to Table 2. The dashed line represents a one-toone correspondence. C1 and B2 (open symbols) served as calibration data.
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6. Conclusions In this work the authors have discussed experimental endurance results of HVQFN-packages soldered onto printed circuit boards. By validating the finite element analyses by a selected series of small, medium and large HVQFN-packages assembled on printed circuit boards, we have tried to determine the board level thermal performance of this family. The discriminating parameters that determine the board level performance of this family are investigated by using analytical and numerical analyses. The following conclusions and recommendations can be formulated: Contrary to the frequently adopted starting point of assuming fixed board properties, they may vary markedly. These variations have a significant impact on the test result which may mask any variations that have been made by design. As theoretically expected, the first order effect of board level reliability performance is dominated by the thermal expansion mismatches. However, there is a strong second order effect that is dominated by bending resistance of the printed circuit board, that is in turn characterized by its thickness and/or Young’s modulus. For the investigated package a thin board with a low stiffness can lead to a factor of 2.5 longer board level life time. It is therefore strongly advised to report on the used board materials when discussing board level reliability results. Also it is likely that the glass transition temperature of the board material affects the solder fatigue life. If Tg enters the range of the temperature cycling test, the mechanism of building up stress changes. To date, this effect can only be addressed qualitatively.
References
Fig. 16. Response surface of normalized fatigue life against: (top) thermal expansion and thickness of board, (bottom) elastic modulus and thickness of board.
son relation (Eq. (1)). In Fig. 16 the results are represented in the form of response surfaces of the lifetime for the two pairs of parameters. Compared to the nominal fatigue life – defined as the lifetime at the following board parameters: thickness 1.6 mm, elastic modulus 17.5 GPa, and expansion coefficient 17.6 ppm/K – the lifetime can vary considerably. If the expansion coefficient of the board increases the lifetime follows, but the more so when thinner panels are used. A factor of 3.5 in lifetime is well possible. With regard to the elastic modulus the effect is to prolong the fatigue life if one would use softer boards, and again: the thinner the board is the better the lifetime will be. As a matter of fact, the elastic modulus has comparatively little influence if the printed circuit boards are thicker than about 2 mm. This same analysis carried out on a center joint leads to essentially the same but far less pronounced results. It will be clear that the numerical model compares better to the experimental results than the analytical attempts. These latter do, however, give a more direct insight in the physics behind the mechanisms that play a role in the fatigue deformation.
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Contents lists available at ScienceDirect
Microelectronics Reliability journal homepage: www.elsevier.com/locate/microrel
Solder-joint reliability of HVQFN-packages subjected to thermal cycling J. de Vries a,*, M. Jansen a, W. van Driel b,c a b c
Philips Applied Technologies, High Tech Campus 7, 5656AE Eindhoven, The Netherlands IMO Back End Innovation, NXP Semiconductors, Gerstweg 2, 6534AE Nijmegen, The Netherlands Delft University of Technology, Mekelweg 2, 2628CD Delft, The Netherlands
a r t i c l e
i n f o
Article history: Received 3 October 2008 Received in revised form 11 December 2008 Available online 4 February 2009
a b s t r a c t In this work experimental results of thermal cycling tests on HVQFN-packages mounted on printed circuit boards are combined with finite element analyses. Validating the finite element analyses by a selected series of small, medium and large HVQFN-packages assembled on printed circuit boards, allows us to determine the performance of this family. To be able to do that, the discriminating parameters that determine the board level performance of this family need to be understood. The emphasis is on the fatigue life of the soldered interconnections as it is influenced by the thermal stress load, the board thickness, and the dimension of the package. Data from different experimental set-ups are compared. An important parameter in this respect is the inclusion of the base material of the panels. The test loads were set to cycling at 40 °C/+125 °C and 20 °C/+100 °C. The results prove that the essential physical properties governing the fatigue life are the stiffness of the complete assembly and the thermal expansion mismatch between the parts. Ó 2008 Elsevier Ltd. All rights reserved.
1. Introduction Over the past few years new types of IC-packages have been developed, among which the quad flat non-leaded (QFN) or micro-lead frame (MLF) packages have become very popular. In this class a variety with a low building height and a heat sink is usually called HVQFN. The advantages of these packages are manifold. They use less board space than other similar packages since they are leadless. Having no solder bumps and leads the electrical properties, such as the self-inductance, are better. The comparatively large heat sink makes them thermally superior. Still, there are some concerns as to the board level reliability. The low solder joint height of typically a few tens of micrometers leads to a much larger shear strain caused by thermal expansion mismatches. And thus the solder fatigue life is reduced. Already some, yet not very much, data have been published on the board level endurance of QFN-type of packages. In one of the earliest papers both the mechanical and the thermal solder joint reliability of bump chip carrier- and QFN-packages was compared [1]. An extensive overview was published that dealt with design parameters of the QFN-package and of the printed circuit board [2]. These comprise such items as the dimension of the solder lands, the height of the interconnection, die size, board thickness, and of course the test condition. One of the first numerical simulation studies that was carried through
* Corresponding author. Tel.: +31 40 27 48765. E-mail address: [email protected] (J. de Vries). 0026-2714/$ - see front matter Ó 2008 Elsevier Ltd. All rights reserved. doi:10.1016/j.microrel.2008.12.007
basically covered the same topics as above [3]. Somewhat similar work combining experiments with modeling was issued in the last few years [4–6]. Yet in all these valuable studies the basic material properties of the printed circuit board were more or less taken for granted. To date one has just begun to address this point [7–9]. After measuring the board material properties the durability of several components – but no QFN – under vibrational or thermal cycling load was simulated. It is the aim of the present work to show the importance to know the board properties. To this end results from various experiments have been evaluated and completed with additional tests. Numerical simulations serve to support the analyses. Whenever possible, analytical models are formulated to illustrate the problem. 2. Experimental issues and simulation For the present work small-sized HVQFN24, medium-sized HVQFN48, and large-sized HVQFN72-type packages were selected as test carrier. In Fig. 1 the layout of such package is shown. They were made with daisy chains to facilitate on-line monitoring of the integrity of the soldered interconnections. In Table 1 the relevant data of the test packages are listed. For all experiments printed circuit boards with four Cu-layers (35 lm) and FR4 were used, the only difference being their layout or thickness. The latter was achieved by adjusting the thickness of the FR4-layer. The panels had NiAu-finish. Each panel could host 30 products.
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In order to capture the actual moment of failure, on-line monitoring of the daisy chain resistance is required. Event detectors serve to detect transient interruptions in the chain. In the present case AnaTech 128/256 STD equipment was used. A failure was defined according to IPC9701 as the first interruption of the daisy chain followed by at least nine repeated events [10]. The threshold value of the daisy chain resistance was set to 1000 X. Commercially available statistical software (Reliasoft, Weibull++ version 6) served to analyze the failure distributions. Weibull or log-normal statistics are commonly used to evaluate these. In the present work Weibull analyses were carried out. The failure distributions are all given with 95% confidence levels; the results are summarized in the lower part of Table 2. In some
Fig. 1. Schematic layout of an HVQFN (left) and back side view on actual sample (right).
The packages were dried for 24 h at 125 °C prior to reflow soldering. Reflow was done at a peak temperature of 250 °C as was determined on the boards. Standard lead-free solder paste (Sn3Ag0.5Cu: SAC305) was applied. For the investigation of the effect of the stand-off, additional solder paste was stencil printed on HVQFN48-packages. The products were reflowed and then the normal assembly process was run. After assembly, the boards were inspected by means of X-ray and cross-sectioning. In Fig. 2 one finds selected pictures of the X-ray inspection of packages immediately after assembling. With some effort even the daisy chains can be discerned. We mention here that the voids in the diepad are typical for the various assemblies studied in this work. In the further analyses voiding was not taken into account. Cross-sections of such packages are shown in Fig. 3. Then the assemblies were subjected to a thermal cycling test in a one chamber system (ESPEC ENX12-7.5 CWL). The test condition was 40 °C/+125 °C with a cycle time of 1 h. In Fig. 4 typical temperature profiles as recorded on the packages are shown, including the profile that was used for the thermal testing at 20 °C/+100 °C/ 1 h-cycle (one chamber Grenco GTTS 125.20S). The temperatures were measured by thermocouples placed inside the soldered joints. Various board level test standards advice to test at least 30 but preferably more products [10]. Table 2 contains the basic elements of the experimental set-up.
Table 1 Package geometry HVQFN. The row ‘‘Experiment ID” refers to Table 2. HVQFN Size (mm) Thickness (mm) Die pad (mm) Die (mm) Coverage (%) Pitch (mm) Experiment ID
24 44 0.85 2.8 2.8 2.0 2.0 51 0.5 B
48 77 0.85 5.3 5.3 2.2 2.2 17 0.5 A, B
48 77 0.85 5.3 5.3 3.5 3.5 44 0.5 C
72 10 10 0.85 6.2 6.2 4.4 4.4 51 0.5 B
Fig. 2. X-ray of assembled packages. Pictures are NOT to scale. (a) HVQFN24. (b) HVQFN48. (c) HVQFN72.
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cases early failures were found, which were suspended from the analyses. Together with the surviving samples these data points are marked with separate symbols in the failure distributions. Numerical simulations were carried out on a quarter model of an assembled package. The moulding compound was simulated with a viscoelastic material model, the solder with a creep model. The finite element suite MARC was used for the modeling. Further details are given in the appropriate sections.
16 14
E' (GPa)
12 10 8 6 4 2 0 0
50
100
150
200
T (°C) Fig. 5. Storage modulus of pcb’s (DMA). See Table 2: A (), B (), C 1.6 mm (—), C 0.8 mm ( ). Fig. 3. Cross–sections of assembled packages. Pictures are to scale; the white scale bar is 3 mm. (a) HVQFN24. (b) HVQFN48. (c) HVQFN72. Half-etched die pad is visible.
length change (ppm)
2000
1500
1000
500
0 0
50
100
150
T (°C) Fig. 4. Temperature profiles 40 °C/+125 °C (solid) and 20 °C/+100 °C (dotted) as measured on the packages.
Fig. 6. Relative length change of pcb’s (TMA). See Table 2: A (), B (), C 1.6 mm (—).
Table 2 Assembly and test parameters: profile, label (ID). Pcb: thickness (t), elastic modulus (E), glass transition temperature (Tg), expansion coefficient (a). Solder joint height (h). Number of samples (N). Weibull-parameters: scale (g) and shape (b); one- (1-p) and two-parameter (2-p) analysis. Open cells: data not determined. HVQFN
24
48
Experiment ID
B1
A
B2
C1
C2
C3
C4
B3
T-profile (°C) t (mm) E (GPa) Tg (°C) a (ppm/K) h (lm) N Analysis type
40/125 1.6 11 174 25.0
40/125 1.6 11 132 13.8 54 60 2-p 3110 4.17
40/125 1.6 11 174 25.0
40/125 1.6 15 131 23.8 28 30 2-p 810 4.17
40/125 0.8 15 125
20/100 1.6 15 131 23.8
20/100 0.8 15 125
40/125 1.6 11 174 25.0
30 2-p 1840 2.98
30 2-p 3090 4.26
30 2-p 7270 6.42
40 2-p 4670 4.01
g b
40 1-p 8400 5.0
72
40 2-p 5520 5.32
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Fig. 7. Set B2: left 100 lm-bar; right 60 lm-bar.
3. Results The elastic modulus and thermal expansion coefficient of complete board stacks including the copper layers were determined. Pieces of the various printed circuit boards were cut out to determine these properties by means of Dynamical Mechanical Analysis (DMA) and Thermal Mechanical Analysis (TMA), respectively. The results of measuring the thermo-mechanical board properties are collected in Figs. 5 and 6 (see also Table 2). Although only FR4-based panels were used in the experiments, there is an appreciable variation in the mechanical properties of the boards. They are all four-layer panels with 35 lm thick Cu-layers. In the boards made from the same material but of different thickness (types C in Table 2) the elastic modulus is about 15 GPa below the glass transition temperature (Tg) which lies around 130 °C. The stiffness of the 1.6 mm pcb is of course higher than for the corresponding 0.8 mm panel. Fig. 5 also shows two other sets of data (A and B): these particular pcb’s are of the same thickness, and both have a lower elastic modulus of about 11 GPa below Tg. However, their thermal expansion coefficients differ considerably, as do their respective glass transition points. All of these differences will undoubtedly have effect on the fatigue life of the assemblies. A higher elastic constant or a larger expansion mismatch must induce more stress during thermal cycling. At several moments during the tests samples were taken out for failure analysis. The failure mode was assessed from cross-sections made of failed samples. In Figs. 7 and 8 some examples clearly reveal solder fatigue. Plastic deformation of the soldered joints and cracks within the solder appear. 4. Discussion of experimental results Per each variation the results will be discussed and compared with literature data whenever these are available. In a few cases
we will use analytical models to illustrate the problem and the advantages of the numerical approach. The technical details of the finite element simulations will be presented in Section 5. The basic elements of the analytical model are given here. For comparing the effects of thermal cycling ranges on the fatigue lifetime (s) one applies the Coffin–Manson relation:
2s ¼ ðDc=2ef Þ1=c ;
ð1Þ
where Dc is the maximum shear strain between the component and the board, ef the fatigue ductility coefficient, and c the ductility exponent respectively. For eutectic SnPb-solder the exponent c is about minus one half, leading to the famous quadratic dependence of the fatigue life on the temperature range. A useful equation to estimate the ductility exponent was given by Engelmaier [11]:
c ¼ c0 þ c1 T center þ c2 lnð1 þ 360=tdw Þ:
ð2Þ
The coefficients ci are specific to the solder material. Further Tcenter is the center temperature of the thermal cycle range, and tdw is the hold time at the extreme temperatures. For three solder alloys the values of ci and the ductility coefficient ef are listed in Table 3 [12,13]. The following equation describes the maximum shear strain (Dc):
Dc ¼ ðDL=hÞDaDT:
ð3Þ
Table 3 Coefficients ci for ductility exponent and ductility coefficient ef for three solder alloys [12,13].
c0 c1 (1/K) c2
ef
SnPb
SAC305
SnAg
0.502 7.34E04 1.45E02 2.25
0.347 1.74E03 7.83E03 3.47
0.416 2.10E-03 1.40E-02 2.25
Fig. 8. Left: set A 100 lm-bar; right: C1 60 lm-bar.
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99.
cumulative failures (%)
90.
50.
10. 5.
1. 100.
1000.
10000.
N-cycles Fig. 9. Weibull distributions. Temperature profile: ( , C1) 40 °C/+125 °C, ( , C3) 20 °C/+100 °C. Pcb 1.6 mm. Labels see Table 2.
L is the span of the soldered joints under consideration, DL their thermal displacement, h is their height, Da is the thermal expansion mismatch between the component and the board, and DT is the temperature difference to which the assembly is subjected. For the fatigue life as function of the solder joint height one has:
s h1=c
ð4Þ
In case of SAC-solder the proposed data in Table 3 lead to a value for the ductility exponent of c = 0.4. 4.1. Variation of test condition
4.2. Effect of solder joint height The influence of the height of the soldered interconnections on the fatigue life is given in Fig. 10. All data were obtained on boards of 1.6 mm thickness (see Table 2, sets A and C1). The smaller standoff of 28 lm leads to a mean time to failure of 810 cycles. By doubling the solder joint height the fatigue lifetime increases to 3110 cycles, which is a factor of 3.8. Other experimental results for comparison could not be found. But one generally acknowledges the beneficial effect of a larger stand-off.
99.
99.
90.
90.
cumulative failures (%)
cumulative failures (%)
In Fig. 9 the Weibull failure distributions of assemblies subjected to two different test conditions are shown. A summary of
the statistical evaluation is listed in the lower half of Table 2 (test runs C1 and C3 with 1.6 mm thick panels). In the first place it must be noted that both distributions run in parallel, as also the shape parameters (b) show. In the second place from the failure analysis – not shown here – the same failure mode was found. This is a strong indication that both tests invoke the same failure mechanism. The milder temperature cycling condition of 20 °C/ +100 °C leads to almost a fourfold longer fatigue life (3090 cycles compared to 810 cycles). In the series with the 0.8 mm thick boards this is the same: 7270 versus 1840 cycles. Astonishingly little reference data for QFN-packages have appeared in the literature dealing with the temperature cycle range. Syed and Kang [2] mention experiments on QFN32-packages of 5 5 mm2 subjected to three different thermal cycling tests. However, only the moment of first failure is reported as the experiments were still running. Cycling at 40 °C/+125 °C/1-h the first failure is observed at 3352 cycles, at 55 °C/+125 °C/40-min this is at 2610 cycles, while after 5813 cycles testing at 0/+100 °C/30min no failures had occurred yet. The same type of data are listed in an application note of Freescale [14]. For QFN48-products of 7 7 mm2 body size cycling at 40 °C/+125 °C/1-h leads to the first failure at 1340 cycles, and at 0/+100 °C/30-min this increases to 3950 cycles. The temperature ranges in our experiments have a ratio of 1.4, which results in an acceleration factor of 2.3. This must be compared to a ratio of about 4 between the experimental lifetimes. Clearly the relatively simple Coffin–Manson approach does not suffice in the current situation.
50.
10. 5.
1. 100.
1000.
10000.
N-cycles Fig. 10. Weibull distributions. Solder joint thickness: ( , C1) 28 lm, ( , A) 55 lm. Pcb 1.6 mm, 40 °C/+125 °C. Labels see Table 2.
50.
10. 5.
1. 100.
1000.
10000.
N-cycles Fig. 11. Weibull distributions. Pcb thickness: ( , C1) 1.6 mm, ( , C2) 0.8 mm. 40 °C/+125 °C. Labels see Table 2.
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99.
was 1254 cycles. In the work of Mercado et al. [4] the relative effect of the panel thickness on the solder fatigue life of BCC32-packages is given. Going from 0.51 to 1.02 mm thickness the lifetime decreases by a factor of 2.8, and from 1.02 to 1.57 mm by a factor of 1.6. In the present case, the assemblies on the thinner printed circuit boards have a fatigue life that is a factor 2.3 longer than the series on the thick boards. Whereas this agrees fairly well with the referenced literature data, an analytical model is much more difficult to give. Straightforward mechanics tells us that the stiffness of a thin slab is proportional to the cube of its thickness. Considering the assembly as a spring system, and assuming that the energy contained in this system is inversely proportional to its lifetime, only a very qualitative description of the effect of board thickness on fatigue life can be obtained.
cumulative failures (%)
90.
50.
10. 5.
4.4. Effect of package size
1. 100.
1000.
10000.
N-cycles Fig. 12. Weibull-distributions. Package size: ( , B3) HVQFN72, ( , B2) HVQFN48, ( , B1) HVQFN24. Pcb 1.6 mm, 40 °C/+125 °C. Labels see Table 2.
From Eq. (4) it follows that the stand-off change from 28 lm to 55 lm should lead to a 5.7-fold increase in fatigue life. However, not only the solder joint height, but also the thermal expansion coefficient of the printed circuit boards used in the two experiments were different. Taking this fact into account the analytical result becomes even worse. 4.3. Effect of stiffness of printed circuit board The stiffness of the printed circuit boards depends on the material constant – the elastic modulus – and the construction – the thickness – of the board. Experimental results for boards of thicknesses ranging from 0.8 to 1.6 mm are shown in Fig. 11 (data sets C1 and C2 of Table 2). Similar experimental results were reported by Syed et al. [2] on QFN72-packages of 10 10 mm2 size subjected to 40 °C/+125 °C cycling. On boards of 1.6 mm thickness the fatigue life was 854 cycles, whereas for 0.8 mm boards this
To investigate the effect of the package body size on the thermal fatigue life, three different HVQFN-types were chosen (see Table 2, series B). However, after 5400 cycles only the two larger packages had a failure rate that allowed for a statistical analysis, the smallest had only four failures. In order to include at least a good estimate for HVQFN24, the width of the distributions (b, Table 2) for the two larger packages was taken as the starting point for a one-parameter Weibull-analysis of the small package: b = 5. In Fig. 12 we show the graphical result. Thus the scale parameters of the three HQVFN-body sizes can be compared. For HVQFN24 the uncertainty is in the range of ±2000 cycles. As for HVQFN72 it must be noted that the lead frame is half-etched which will increase the compliance of the package and hence the lifetime is expected to be longer. Tee et al. [3] report on board level solder joint reliability of QFNpackages with a different outline. These were tested by cycling at 40 °C/+125 °C with a cycle period of 40 min. The larger QFN52 of 8 8 mm2 has a fatigue life in the range of 631–1426 cycles. For the smaller QFN20 of 4 4 mm2 this is 2743–4894 cycles. The ranges depend on design variations. 5. Numerical model The discussions of the experimental results prove the shortcomings of analytical calculations to analyze the results. In this section
Fig. 13. Quarter FE-model (left), die pad (top right), cross-section showing also die and lead frame (bottom right).
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150
1.E-03
8.E-04
6.E-04 50
T (°C)
strain (-)
100
4.E-04
0 2.E-04
-50
0.E+00 0
2000
4000
6000
8000
time (s) ), and plastic strain (left axis,
we will briefly discuss the finite element model and the results of using it to a parameter sensitivity analysis. The creep model proposed by Darveaux was used to simulate the behavior of the solder material [15]:
n dccr GðTÞ s eQ =kT : sinh h ¼C T GðTÞ dt
ð5Þ
Here dc/dt is the creep strain rate, G is the shear modulus, h the stress level where the power law breaks down, s the applied stress, n the stress exponent, and Q is the activation energy. For Sn3.5Ag the following values were taken: C = 0.4539, G(T) = 19,300 69(T 273) [MPa], h = 1500, n = 5.5, and Q = 0.5 eV [15], see also [16,17]. The finite element model consists of a quarter of the complete assembly (see Fig. 13). Two thermal cycles have been simulated (see Fig. 14). The strain was determined by averaging over one complete soldered connection at the corner of the package since this is the most critical joint. In Fig. 14 the plastic strain and the creep strain are shown as they build up during thermal cycling. Plastic strain is only 10% of the creep strain. As the temperature is at its lower extreme, the creep mechanism is hardly active, while at the higher extreme only little creep occurs because the stress is much lower. Only during the temperature ramps an appreciable amount of creep strain develops. Application of the finite element model to various assemblies leads to a satisfactory result as depicted in Fig. 15. Here, one finds the lifetime simulated for the B-and C-series and calibrated against the respective experimental results of HVQFN48 at 40 °C/+125 °C on 1.6 mm boards (see Table 2, series C1 and B2). One recognizes the outcome of testing at a different cycling condition which was discussed in Section 4.1 (Table 2, series C3 and Fig. 9). Likewise, the tests carried out on packages soldered onto a thinner board, discussed in Section 4.3, are shown (Table 2, series C2 and Fig. 11). To further validate the model, also other HVQFN-packages were modeled and compared to experimental test results. These are the smaller HVQFN24 (4 4 mm2) and the larger HVQFN72 (10 10 mm2) which can be found in Table 2 series B1, B3 and Fig. 12. Finally, the test series with a larger solder height (Table 2, series A) was simulated. As one can see in Fig. 15 this result does not
). (For interpretation of the references in color in this figure legend,
compare very well to the experiment. The main difference is the much lower thermal expansion coefficient of the printed circuit board used for this experiment (see Fig. 6 and Table 2). Still, keeping this in mind, one may safely conclude that the numerical model is sufficiently reliable to proceed with the sensitivity analysis. The analyses of the experiments sofar lead to the notion that the thickness, the elastic modulus, and the thermal expansion coefficient of the printed circuit boards govern the fatigue life of the interconnections and variations therein should not be neglected. In this case one thermal cycle was modeled. Three board parameters were varied (in parentheses the nominal value): – Thickness 0.5–2.5 mm (1.6 mm). – Elastic modulus 10–20 GPa (17.5 GPa). – Thermal expansion coefficient 10–25 ppm/K (17.6 ppm/K). Again, one of the corner joints was used to average the strain that was translated to fatigue lifetime by means of the Coffin–Man8000
simulation (cycles)
Fig. 14. Temperature profile (right axis, red line), creep strain (left axis, the reader is referred to the web version of this article.)
B2
6000
B1
4000
B3 C3
2000
C1 C2
A
0 0
2000
4000
6000
8000
10000
experiment (cycles) Fig. 15. Simulated lifetime (63%-failures) against experimental data (with 95% confidence levels). Data labels refer to Table 2. The dashed line represents a one-toone correspondence. C1 and B2 (open symbols) served as calibration data.
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6. Conclusions In this work the authors have discussed experimental endurance results of HVQFN-packages soldered onto printed circuit boards. By validating the finite element analyses by a selected series of small, medium and large HVQFN-packages assembled on printed circuit boards, we have tried to determine the board level thermal performance of this family. The discriminating parameters that determine the board level performance of this family are investigated by using analytical and numerical analyses. The following conclusions and recommendations can be formulated: Contrary to the frequently adopted starting point of assuming fixed board properties, they may vary markedly. These variations have a significant impact on the test result which may mask any variations that have been made by design. As theoretically expected, the first order effect of board level reliability performance is dominated by the thermal expansion mismatches. However, there is a strong second order effect that is dominated by bending resistance of the printed circuit board, that is in turn characterized by its thickness and/or Young’s modulus. For the investigated package a thin board with a low stiffness can lead to a factor of 2.5 longer board level life time. It is therefore strongly advised to report on the used board materials when discussing board level reliability results. Also it is likely that the glass transition temperature of the board material affects the solder fatigue life. If Tg enters the range of the temperature cycling test, the mechanism of building up stress changes. To date, this effect can only be addressed qualitatively.
References
Fig. 16. Response surface of normalized fatigue life against: (top) thermal expansion and thickness of board, (bottom) elastic modulus and thickness of board.
son relation (Eq. (1)). In Fig. 16 the results are represented in the form of response surfaces of the lifetime for the two pairs of parameters. Compared to the nominal fatigue life – defined as the lifetime at the following board parameters: thickness 1.6 mm, elastic modulus 17.5 GPa, and expansion coefficient 17.6 ppm/K – the lifetime can vary considerably. If the expansion coefficient of the board increases the lifetime follows, but the more so when thinner panels are used. A factor of 3.5 in lifetime is well possible. With regard to the elastic modulus the effect is to prolong the fatigue life if one would use softer boards, and again: the thinner the board is the better the lifetime will be. As a matter of fact, the elastic modulus has comparatively little influence if the printed circuit boards are thicker than about 2 mm. This same analysis carried out on a center joint leads to essentially the same but far less pronounced results. It will be clear that the numerical model compares better to the experimental results than the analytical attempts. These latter do, however, give a more direct insight in the physics behind the mechanisms that play a role in the fatigue deformation.
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