**Transient Thermal and Stress Analysis for TSOP Devices**

Mark S. Fan

NASA Technology Validation Assurance Program

Goddard Space Flight Center

Greenbelt, Maryland 20771

301-902-4357

fan@package.gsfc.nasa.gov

Gregory Rose

Electronic Packaging and Processes Branch

Goddard Space Flight Center

Greenbelt, Maryland 20771

301-286-4280

glrose@pop300.gsfc.nasa.gov

**Abstract**

*Plastic Encapsulated Microcircuits (PEMs) are now being used in NASA
spaceborne electronic systems mainly for their cost effectiveness. Among
those PEMs, thin-small-outline-plastic (TSOP) surface mount devices are
of particular interest because of certain benefits associated with light
weight and compact volume. In a typical solder reflow process, a TSOP
device is subject to a thermal transition from room temperature to 260
°C, and multiple sites of interfacial delamination and cracking have
been observed in the plastic and I/O lead interface. This could be
critical since moisture diffused through the cracks could cause
corrosion in the metallization. The purpose of this study is to
understand the mechanical impact of that thermal transition on the TSOP
device through transient thermal and stress analyses. It has been
revealed that the TSOP device generally has a good heat conduction
capacity because of the thin encapsulation. However, excessive shear
stresses are induced in the plastic-lead and plastic-die paddle
interfaces, which could cause delamination and cracking as have been
found during the inspection of reflowed TSOP devices. Due to the complex
nature of this problem, a more accurate simulation will need to include
plastic creep and nonlinear time-dependent fracture models*.

Key words: PEMs, TSOP, FEA, surface mount solder joint, heat conduction,
thermal stress, delamination.

**1. Introduction**

NASA GSFC is currently evaluating TSOP surface mount devices for
spaceflight applications. Typical configurations of a 16 MB memory TSOP
device are shown in Fig. 1-4 which have been used in the analysis.
Although there are many different molding compounds on the market, only
one has been selected for this study and it is used by at least one
major semiconductor company in USA. This plastic material is labeled
Compound X in this paper. Many advantages are associated with the use of
the TSOP SMT packaged devices in spaceborne electronic systems, such as
compact volume, light weight, low profile PWBs, better vibrational
resistance and more efficient heat removal from the chip. In this study,
transient temperature and thermal stress distributions during thermal
shock from room temperature up to 260 °C are the key interests since the
shock response of this ultra-thin package is expected to be quite
different from that of the conventional packages. These transient
physical processes are modeled with 3-D finite element techniques as
part of the evaluation effort.

This work is mainly focused on heat transfer characteristics as well as
thermal stresses caused by mismatches between different materials, and
assessment of possible damage. Although popcorning is a popular
phenomenon in recent studies, it is assumed here that the devices are
well baked and dried before the soldering operation \[1\].

![](media/image1.emf){width="2.30625in" height="1.5006944444444446in"}

**Fig. 1. Top view of the TSOP**

![](media/image2.emf){width="2.3875in" height="1.525in"}

**Fig. 2. X-ray top view of the TSOP**

![](media/image3.emf){width="2.7423611111111112in"
height="0.5590277777777778in"}

**Fig. 3. X-ray side view of the TSOP**

Paddle: 6

Die: 14 38

Lead: 6

9

**Fig. 4. Internal dimensions (not to scale)**

**(Units: mil. 1.0 mil=25.4 mm)**

**2. Material Properties**

It is well known that thermal and structural properties of epoxy molding
compounds are highly temperature-dependent. For Compound X, the
manufacturer only provides very limited property data over the
temperature range. FEA modeling results would be unrealistic without
using property data covering the full temperature range. In order to
fill the gap, GSFC conducted a series of detailed measurements over a
large temperature range. Due to measurement hardware limitations,
thermal conductivity of the compound was measured only at 22 °C and 50
°C. Fig. 5 shows stress-strain curves from 22 °C to 166 °C, and Fig. 6
gives temperature-dependent failure limit of the stress from 22 °C to
250 °C. In Fig. 7, tensile and flexural moduli are plotted as a function
of temperature from 22 °C to 166 °C. The coefficient of thermal
expansion (CTE) is shown in Fig. 8 as measured from 20 °C to 260 °C.
Temperature-dependent Poisson's ratio is plotted in Fig. 9. Table 1
lists other property data used in the analyses.

![](media/image4.emf){width="2.422222222222222in"
height="1.8708333333333333in"}

**Fig. 5. Stress-strain curves of the plastic**

![](media/image5.emf){width="3.1083333333333334in"
height="2.342361111111111in"} **Fig. 6. Stress failure limits of the
plastic**

![](media/image6.emf){width="2.6659722222222224in"
height="2.2284722222222224in"}

**Fig. 7. Moduli of the plastic**

![](media/image7.emf){width="2.7430555555555554in"
height="1.8833333333333333in"}

**Fig. 8. CTE of the plastic**

![](media/image8.emf){width="2.747916666666667in"
height="1.6513888888888888in"}

**Fig. 9. Poisson's ratio of the plastic**

+--------------------+----------+------------------+------------------+
| [Mate              | [Si]{.un | [Ni/Fe           | [Mold            |
| rials]{.underline} | derline} | Al               | ing]{.underline} |
|                    |          | loy]{.underline} |                  |
+--------------------+----------+------------------+------------------+
| Young's Modulus    | 120      | 145              | See Fig. 6       |
|                    |          |                  |                  |
| (GPa)              |          |                  |                  |
+--------------------+----------+------------------+------------------+
| Poisson's          | 0.28     | 0.31             | See Fig. 8       |
|                    |          |                  |                  |
| Ratio              |          |                  |                  |
+--------------------+----------+------------------+------------------+
| Density (kg/m^3^)  | 2330     | 8110             | \-               |
+--------------------+----------+------------------+------------------+
| Specific Gravity   | \-       | \-               | 2.014            |
+--------------------+----------+------------------+------------------+
| Thermal            | 148      | 129              | 22 °C: 0.86      |
| Conductivity       |          |                  |                  |
|                    |          |                  | 50 °C: 0.89      |
| (W/m°C)            |          |                  |                  |
+--------------------+----------+------------------+------------------+
| Specific Heat      | 712      | 477              | 882              |
| (J/kg°C)           |          |                  |                  |
+--------------------+----------+------------------+------------------+
| CTE                | 2.6      | 4.45             | See Fig. 7       |
|                    |          |                  |                  |
| (ppm/°C)           |          |                  |                  |
+--------------------+----------+------------------+------------------+

**Table 1. Material properties for analysis**

**3. FEA Model and Constraints**

Because of symmetry, a one-quarter FEA model is created for the TSOP
device (Fig. 10). Two types of analysis are conducted: transient thermal
shock from 22 °C to 260 °C, and thermal stress induced by thermal shock.

In the thermal shock analysis, the initial temperature of the TSOP
device is 22 °C. The environmental temperature surrounding the device is
increased to 260 °C using a temperature step function at the time t = 0
sec (Fig. 11). Transient heat conduction starts from that moment and the
heat flux moves from the skin towards the center of the device till
thermal equilibrium is reached.

Of particular interest in thermal stress analysis is the maximum stress
distribution after thermal equilibrium is established. In thermal stress
analysis, two symmetric displacement conditions are incorporated along
the two symmetric planes of intersection. In addition, the end of each
metal lead is constrained to prevent rigid body movement.

![](media/image9.emf){width="2.442361111111111in"
height="1.5222222222222221in"}

**Fig. 10. One-quarter TSOP FEA model**

![](media/image10.emf){width="2.2979166666666666in" height="1.6in"}

**Fig. 11. Thermal shock profile**

**4. FEA Results**

Fig. 12 shows the transient heat conduction process in the silicon chip.
After 0.1 sec from the thermal shock, the highest temperature on the
chip surface is increased by approximately 55 °C, by 3.4 sec thermal
equilibrium is essentially reached. The temperature gradient across the
chip is initially small due to constant room temperature (22 °C), but
with more and more heat flux, this gradient peaks dramatically and
disappears quickly because of rapid thermal equilibrium. Fig. 13 shows
this change in temperature gradient.

![](media/image11.emf){width="2.1868055555555554in"
height="1.7944444444444445in"}

**Fig. 12. Chip transient temperature rise**

![](media/image12.emf){width="2.071527777777778in"
height="1.7368055555555555in"}

**Fig. 13. Transient temperature gradient change in the chip**

For the overall TSOP device, the time needed to reach thermal
equilibrium is longer than 3.4 sec, because of the differences in
thermal diffusivity of the materials. Fig. 14 illustrates this process,
and it can be seen that the practical equilibrium time is about 6.6 sec.

![](media/image13.emf){width="2.1958333333333333in"
height="1.8402777777777777in"}

**Fig. 14. TSOP transient temperature rise**

The distribution of maximum principal stress in the silicon chip at 260
°C is obtained by the transient stress analysis. The peak stress,
occurring around the corner of the chip, is 198 MPa. Compared to the
averaged fracture strength of 1.0 GPa, this shock-induced stress is well
within the safe region. Similarly, the distribution of the maximum shear
stress in the plastic molding compound shows a peak value of 66 - 68 MPa
occurring at the interface of the plastic and the die paddle, as well as
the plastic and the I/O lead. These are considered the likeliest sites
for possible interfacial delamination under this thermal shock because
the lab shear testing shows that the compound's failure shear stress
(15 - 65 MPa) lies below the actual peak stress in those regions.

Using thermal stress level as a delamination criterion may be inadequate
for mainly two reasons. First, mechanical stresses induced by moisture
expansion inside the package at high temperature (which is not the
concern in this study because of sufficient baking before reflow) could
cause more significant damages than thermal stresses during heating up
stage in a reflow \[2\], and a simultaneous solution using FEA for both
thermal stress and pressure-induced stress is difficult. Often the total
stress effect is underestimated and only thermal stress is taken into
consideration. This combined stress issue remains largely unsolved,
although evaluating the effect of the two stress fields separately is
straightforward. Secondly, due to r^-1/2^ type stress singularity, local
stresses around a sharp corner or a crack tip are non-deterministic.
Recently several attempts have been made to analyze plastic package
delamination/cracking problem with classical linear elastic fracture
mechanics \[3\], or elastic-plastic fracture mechanics (\[2\], \[4\] and
\[5\]). Through the use of stress intensity factors and J contour
integral, the results of these work provided viable means dealing with
internal delamination of plastic packages.

However, there still exist fundamentally unsolved issues in this
critical area. First, all the work mentioned above assumed a
through-thickness crack mechanism, i.e, plane stress or plane strain.
The inspection of reflowed TSOPs revealed that typical
delamination/cracking sites were neither plane stress nor plane strain.
Secondly and most importantly, fracture behavior of thermoset plastics
is quite different from that of metals and other brittle materials, and
applying fracture mechanics concepts for metals to plastics may be
misleading. A salient feature of thermoset plastic material is the
rate-dependent viscoelastic behavior during deformation. Plastics do
yield under large strain with a mechanism of shear deformation.
Typically, a thermoplastic goes through state transitions when
temperature changes significantly. When temperature is close to its
glass transition point (T~g~), the polymer exhibits a "leathery"
behavior. At higher temperature, it is in rubbery state. If temperature
reaches a certain high level, the polymer will be unable to carry any
load and will behave like a viscous fluid. For Compound X, the measured
T~g~ is about 125 °C.

To account for the time-dependent viscoelastic behavior, two quantities
must be included in the analysis: relaxation modulus E(t) and creep
compliance D(t), which can be measured by fixing strain and stress
respectively.

> E(t) = s(t)/e~0~ (1)
>
> D(t) = e(t)/s~0~ (2)

Modeling thermal shock of this magnitude could be considered too harsh
and unrealistic for practical purposes such as solder reflow process.
Realistically, the temperature profile in this reflow process behaves
more like a thermal cycle with a gradual change from 22 °C to 260 °C.
The difference in the resultant stresses in the compound with a thermal
cycle is mainly due to creep, a stress relaxation process. Therefore
stresses induced at 260 °C under thermal cycling will be significantly
lower. Considering the large shear failure range and the lower bound (15
MPa) for this compound, it is possible that delamination could still
occur in those weak interfaces even with creep in the compound. At
present, several material-dependent coefficients and exponents in the
constitutive model for creep are not available for this compound. A more
extensive material testing effort is required in order to obtain those
properties.

Let us compare stress change in the chip under the shock condition with
that under a gradual temperature change, the latter being closer to
actual solder reflow profile. Fig. 15 shows a hypothetical profile which
has a ramp time of 5 minutes. The changes of maximum principal stress
under the shock profile and the gradual profile are plotted in Fig. 16
and 17 respectively. Without creep effect in the model, resultant peak
stresses (198 MPa or 28700 psi) are the same, but the rate of increase
is different.

![](media/image14.emf){width="2.2597222222222224in"
height="1.7965277777777777in"}

**Fig. 15. Hypothetical reflow profile**

![](media/image15.emf){width="2.279861111111111in"
height="1.9583333333333333in"}

**Fig. 16. Transient max. principal stress in the chip under shock
profile**

![](media/image16.emf){width="2.301388888888889in"
height="1.9965277777777777in"}

**Fig. 17. Transient max. principal stress in the chip under "reflow"
profile**

**5. Conclusions**

Finite element transient thermal analysis and thermally induced stress
analysis are performed for a TSOP device made with Compound X. It is
found that this thin plastic package takes only a few seconds to reach
thermal equilibrium in a thermal shock environment from 22 °C to 260 °C.
This indicates that this type of device has a good heat conduction
capacity.

Several interfacial areas in the device have excessive shear stresses
and are susceptible to interfacial delamination under the thermal shock
conditions. Due to plastic creep (that has not been taken into account),
the possibility of delamination is reduced but not entirely eliminated.
A more complete model should account for the viscoelastic behavior of
the compound, but this requires far more testing that is ongoing.
Analysis incorporating nonlinear creep process, together with 3-D
fracture mechanics determining delamination/cracking growth and arrest,
will constitute our future endeavor in the TSOP reliability evaluation.

**6. Acknowledgment**

The authors would like to thank Ann Garrison-Darrin (GSFC code 310) and
Jack Shaw (GSFC code 312) for their generous support, and Dr. Henning
Leidecker (GSFC code 311) for many useful discussions on behavior of the
molding compound material. The testing of Compound X by GSFC code 313
personnel, including Michael Viens, Jackie Townsend, Chuck Powers and T.
Van Sant, is also deeply appreciated.

**7. References**

\[1\] T. Steiner and D. Suhl, "Investigations of large PLCC package
crack during surface mount exposure," *IEEE Trans. Compon., Hybrids and
Mfg. Tech*., Vol. 10, No. 2, pp. 209-216, 1987.

\[2\] H. Lee and Y. Y. Earmme, "A fracture mechanics analysis of the
effects of material properties and geometries of components on various
types of package cracks*," IEEE Trans. Compon., Pkg. and Mfg. Tech*.,
Vol. 19, No. 2, pp. 168-178, 1996.

\[3\] T. Saitoh, "Numerical stress analysis of resin cracking in LSI
plastic packages under temperature cycling loading," *IEEE Trans.
Compon., Pkg. and Mfg. Tech*., Vol. 19, No. 3, pp. 593-600, 1996.

\[4\] J. Sauber, *et al*., "Fracture properties of molding compound
materials for IC plastic packaging," *IEEE Trans. Compon., Pkg. and Mfg.
Tech*., Vol. 17, No. 4, pp. 533-541, 1994.

\[5\] A. A. O. Tay, *et al*., "Predicting delamination in plastic IC
packages and determining suitable mold compound properties," *IEEE
Trans. Compon., Pkg. and Mfg. Tech*., Vol. 17, No. 2, pp. 201-208, 1994.