September 2015, Friedrichshafen, Germany
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European Microelectronics
Packaging Conference
Ni–Sn Solid-Liquid Interdiffusion (SLID) Bonding for Thermo-Electric
Elements in Extreme Environments – FEA of the joint stress
Andreas Larsson*1,2, Torleif A. Tollefsen3, Ole Martin Løvvik4 and Knut E. Aasmundtveit2
TECHNI AS, Borre, Norway
HBV – Buskerud and Vestfold University College, Borre, Norway
3
TEGma AS, Kristiansand, Norway
4
SINTEF Materials and Chemistry, Oslo, Norway
1
2
* Corresponding Author: andreas.larsson@hbv.no, +47 473 60 374
Abstract
Ni–Sn solid-liquid interdiffusion (SLID) bonding was investigated for use in extreme thermal conditions. Energy
harvesting by thermoelectrics push for utilization of ever increased temperature gradients to improve energy
conversion efficiency. Exposure to large temperature gradients induce thermomechanical stress in joints that may lead
to catastrophic device failure by fractures in the joint. Finite element analysis of skutterudite CoSb3 joined by Ni–Sn
SLID bonding to alumina substrates was performed. The bond structure was CoSb3 / TiN / Ni / Ni3Sn4 / Ni / Cu / Al2O3.
Temperature gradients of up to ~100 °C/mm at temperatures up to 500 °C were employed. Two types of models were
compared; (1) one element bonded to a substrate on one side and (2) one element symmetrically bonded to substrates
on both sides. The results show that the stress field is dominated by the residual stress from the process, with limited
contributions from external loads and system configuration.
1
Introduction
Thermoelectric generators (TEG) convert heat energy to
electrical energy when exposed to a thermal gradient
across the material. Negatively and positively doped
semiconductors are placed alternately in series with an
electrical load to form a circuit. A temperature gradient is
applied in parallel across the thermoelectric materials,
typically called elements or sometimes legs. The high
temperature causes electrons (or holes) to break free from
the atoms on the hot side by a shift in the Fermi level.
This create a particle flow from the hot side of the
material to the cold side, i.e. generating a current. High
temperature (HT) applications is particularly interesting
due the possibility of large temperature gradients across
the elements. HT applications suitable for waste heat
regeneration by TEGs include hot jet engines [1], exhaust
systems for combustion engines [2], and large scale
industrial applications [3].
One of the toughest challenges to make reliable high
temperature TEGs is to develop a compatible
interconnect. Temperatures may extend up to several
hundreds of degrees Celsius combined with extreme
temperature gradients beyond 100 °C/mm. Additional
application specific requirements, such as a large number
of thermal cycles [2], make this challenge particularly
difficult. At high temperatures diffusion processes, as
well as thermo-mechanical stress activated mechanisms
such as fatigue and creep, are accelerated. Joints
comprising dissimilar materials that is exposed to
ISBN 978-0-9568086-1-5
temperature changes or thermal gradients may induce
high stresses in the joints which could ultimately lead to
a catastrophic fracture [4][5][6]. Mo–Cu bonded to the
TE material CoSb3 (skutterudite) for high temperature
joints have been investigated with promising results. The
joint showed a residual shear strength capacity of more
than 10 MPa after exposure to 550 °C for 480 hours [7].
Mo–Cu may be engineered to have a suitable coefficient
of thermal expansion (CTE) to match the skutterudite,
but it is expensive. TEG applications may often be cost
sensitive due to large fabrication volumes. Thus, Cu
traces on Alumina substrates may be preferred. Although,
the significant CTE mismatch of Cu, skutterudite, and
Alumina create significant thermomechanical stress
fields at large temperature differences. Reliable high
temperature joining technologies solders are also sparse
when the temperature reach 500 °C or beyond [8].
Solid-Liquid Interdiffusion (SLID), sometimes referred
to as Transient Liquid Phase (TLP), bonding is a
promising technique that have gained a strong interest in
recent years [9][10]. In particular, the technology have
been explored for high temperature applications since the
operation temperature may be significantly higher than
the process temperature [11]. The SLID process utilize
interdiffusion between two components in a bi-metal
system. The interdiffusion creates a phase transformation
changing material properties, such as the melting
temperature. One interesting binary material system is
the Ni–Sn system. It have a CTE between that of Cu and
CoSb3, and Alumina. It is cost effective compared with
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September 2015, Friedrichshafen, Germany
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other common SLID systems with potential for
utilization in applications at 500 °C and beyond [12]. It
is rather unexplored as a SLID system, but has been
explored for MEMS packages [13]. Its phase diagram is
shown in Figure 1. The SLID process can create joints
comprising the intermetallic phases; Ni3Sn, Ni3Sn2 and
Ni3Sn4, which all have melting temperatures about
800 °C or more [14]. The Sn rich phase, Ni3Sn4, is known
from processes comprising Sn based solders on pads with
an electroless nickel immersion gold (ENIG) finish. The
other two intermetallic phases are not yet thoroughly
explored.
expected that all the three intermetallic phases have
similar thermomechanical properties, and that Ni3Sn4
will be the dominant phase by volume in the final joint.
Material properties were modelled as temperature
dependent, when possible. Ni, Cu and TiN were modelled
as elastic-plastic materials with an isotropic hardening
coefficient,  t , or as perfectly plastic. The von-Mises
yield criterion and material yield limit,  y , was used for
the elastic-plastic material models. The other materials
were modelled as pure elastic materials. Utilization
factors, UFe, p , were derived from the equivalent stress,
 e , and the effective plastic strain, ep , normalized by
the material strength,  b , and maximum elongation,
 max , according to (1) and (2).
e
b
(1)
 ep
 max
(2).
UFe 
UF p 
Fabricating a TEG may require bonding of the
thermoelectric element to the two substrates sequentially.
Thus, two basic configurations of the joint stress were
evaluated.
1.
2.
Figure 1. The Ni–Sn phase diagram by [15].
This paper investigates the stress field inside a proposed
Ni–Sn joint in a CoSb3 based TEG. Firstly, its stress
distribution at room temperature after bonding to a first
substrate is analyzed. Secondly, its stress distribution at
various elevated temperatures after bonding is studied.
Bond layer structure dimensions may have an impact on
the stress state [16], thus a parametric study of bond layer
thickness impact on the joint stress state is also presented.
Lastly, impact of a second symmetric joint and substrate
on the hot side is analyzed.
2
Methods and materials
The material configuration was CoSb3 / TiN / Ni / Ni3Sn4
/ Ni / Cu / Al2O3 (Figure 2). The characteristic properties
of the materials used in the model are shown in Table 1.
Data were extracted from pertinent literature. The
thermoelectric element was modelled as CoSb3 with a
TiN adhesion layer. Prior bonding, Sn is sandwiched
between Ni seed layers. During bonding, interdiffusion
between Ni and Sn form a symmetric Ni / Ni3Sn / Ni3Sn2
/ Ni3Sn4 structure. Although all three intermetallic
compounds along with Ni and Sn in mixtures or solid
solutions may exists in the final joint [17], the joint was
modelled as a pure Ni3Sn4 between Ni layers. It was
ISBN 978-0-9568086-1-5
Single sided model with one substrate on the
cold side – Simulating the first bonding process
to the first substrate.
Double sided model with a second substrate on
the hot side as well. – Simulating the second
bonding process to the first substrate.
The stress state was analyzed for one single element
joined to the substrates. Note that real elements in a
module may in addition be loaded by a global load caused
by eccentricity of element location with respect to the
global neutral point of the module [16].
Typical element and substrate dimensions in a TEG were
used. The thermoelectric element height was 5.0 mm
with a square foot print of 5.0·5.0 mm2 and the substrates
were 635 µm thick. The bond layer thicknesses, d , of
the Ni and Ni–Sn layers were varied between 1 µm and
20 µm. A fixed temperature of 25 °C was applied on the
cold side. The temperature on the hot side was varied
between 25 °C and 500 °C. The strain reference
temperature of the joint materials was set equal to the
process temperature; 350 °C. The system was simulated
at thermal equilibrium. Two symmetry planes were used;
one along a diagonal and one at a mid-plane of the
element, i.e. ⅛ was modelled. The models were simply
supported and the mesh was found adequate. Results
were extracted from regions sufficiently far away from
FEM induced singularities, such as edges along
dissimilar materials (cf. [18][19]). An illustration of the
modelled conditions is shown in Figure 3.
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Table 1: Characteristic material properties and dimensions.
(W/mK)
(ppm/K)
(GPa)
(1)
(MPa)
(MPa)
(GPa)
t
 max
2.6-2.91
3500
23.5-28.81
0.7
65-881
1-10
19.6
2-20
373-3981
10
22
635
1
Modeled with temperature dependent property
2
Perfectly plastic
10
8.8-9.41
13.4-16.61
13.9-211
16.7-20.31
8
140
250
180-2251
133
106-1281
300
0.23
0.34
0.29-0.321
0.3
0.34-0.361
0.23
2500
57-961
13-481
-
385
370-1801
55-2211
700
02
0.5
0.2
-
25
51-731
25-551
-
Component / Material
Element / CoSb3
Adhesion layer / TiN
Seed layer / Ni
Bond line / Ni3Sn4
Adhesion layer / Cu
Substrate / Al2O3 96%
d
(µm)



E
y
b
(%)
the substrate compress the edges around the joint for all
studied temperature gradients (cf. the compressive peal
stress in Figure 8 below). At higher temperatures the
backside of the substrate warps in a S-shape mode. This
is caused by the dissimilar material properties, Young’s
modulus, E , Poisson’s ratio,  , coefficient of thermal
expansion,  , combined with the applied temperature
gradient and Tref. The metals in the joint have a
significantly larger  than both the ceramic substrate and
the thermoelectric element resulting in the warpage
around the joint periphery forming the S-shape.
All evaluated configurations of the model show similar
generic characteristics regarding type/direction of stress
components, deformation etc. The absolute relative stress
level between the components varies between
configurations.
In general, the stress state is typically low (Figure 6) as a
result of the fine global thermomechanical match of the
CoSb3 and Al2O3 in the system ( E, and  ). However,
Figure 7 and Figure 8 illustrates a significant stress level
in the joint’s intermetallic layer and that the stress is
tensile. Although, the peel stress (  z ) is compressive
near the joint periphery. The other joint layers show
significantly lower stress levels (Figure 9). This indicate
that void nucleation and fractures could be expected to
initiate from the Ni3Sn4 layer by Mode I loading. Cracks
originating from the element may also be envisaged, if
surface finished is not cared for during dicing or if the
sintered material is imperfect with microvoids etc. Note
that, it is not possible to conclude whether the joint will
fail or not since no strength data of the Ni3Sn4 phase have
been found in the pertinent literature. It is also
noteworthy that the Cartesian stress components coincide
well with the principal stresses (not illustrated explicitly
here), i.e. 1   2   x   y   3   z , which indicate
a stress state with limited shear stress (cf. Figure 8): The
TiN layer remains in the elastic region for all evaluated
configurations. Both the Ni and Cu layer is plastically
strained. The effective plastic strain is typically less than
0.3% and 1% in the layers, respectively. Thus, long term
effects caused by strong stress fields, cyclic loading or a
combination, e.g. fatigue failures from thermal cycling,
or creep failures must be considered. Time and history
dependent changes such as hardening (kinematic or
isotropic), microstructural changes, e.g. grain growth, or
stress relaxation by creep will change the sequential joint
stress state, and needs to be analyzed appropriately in
subsequent studies.
The models show warpage to be approximately 10-20 μm
(Figure 4 and Figure 5) at room temperature. Warpage of
The joint stress state was lowest at room temperature
(Figure 10), i.e. without an applied thermal gradient
Figure 2. Illustration of materials configuration and
main dimensions used in the single sided models.
Figure 3. Illustration of application of the used boundary
conditions.
3
Results and discussion
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across the device. Heating the device increases the stress
levels by less than 15%. Thus, the residual stress from
fabrication dominates the joint stress state for
temperature gradients up to at least 100 °C/mm.
Varying the layer thickness had little effect on the stress
levels in the system (Figure 11). This indicates that the
joint was too stiff to significantly benefit from optimizing
the relative thickness of joint layers for layer thicknesses
of 1 μm or more.
Adding another symmetric substrate and joint to the hot
side of the element did not affect the stress distribution
significantly (Figure 12). The principal in-plane stress
components still dominate the stress field. A small
reduction in the equivalent stress near the center of the
joint is observed. This is due to the cancelation of the
moments from the geometrical symmetry between
components with a mismatch in expansion coefficients,
where CoSb3 and Al2O3 are the main contributors to the
effect.
Joint center
Joint corner
Figure 6. Joint utilization (%) at 375 °C. The utilization
is around 10-15%% for the element near the joint, and
reach around 40% at the corner. Utilization of the
substrate is less than 20%. The joint layers typically show
a utilization of a few percent.
Figure 7. Equivalent stress field in the Ni3Sn4 layer near
the joint corner at Thot = 375 °C. At 1.5 mesh elements
away from the hot spot (about 5 times the layer thickness)
the stress level is less than 350 MPa (cf. Figure 8).
Figure 4. Total displacement in μm (x50) of a model at
room temperature after bonding, illustrating a
compressive deformation around joint borders. Black
lines show the undeformed geometry.
Figure 8. Stress components in the Ni3Sn4 joint layer from
the joint center out to the joint corner (cf. Figure 4). The
equivalent stress is mainly dominated by the in-plane
normal stress components (x & y). Near the corner (>2.5
mm) the out-of-plane component (z) grows rapidly.
Figure 5. Deformation in μm (x50) at Thot = 25 °C, 250 °C
and 500 °C (left to right). Black lines show the
undeformed geometry. A S-shape of the substrate is
clearly visible at 500 °C.
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Location of
joint corner
Figure 9. Equivalent stress in joint layers, including the
interfaces at element and substrate (excluding the Ni3Sn4
layer which is presented in other figures).
Figure 12. Stress in the Ni3Sn4 joint layer for two similar
models at room temperature. Both models show similar
characteristics. Note the small drop in stress level near
the center (<2 mm) originating from the symmetric
moment by a symmetric expansion mismatch of the
substrates and element.
4
Conclusions
This study show that a Ni–Sn SLID joint in a CoSb3
based thermoelectric generator cause high stress levels in
the intermetallic layer in the joint and that adjacent layers
are plastically deformed. Thus, failure by fatigue, creep
or fracture may arise in real joints. Experimental
characterization
of
joint
phases
and
their
thermomechanical properties needs to be added to the
study before final conclusions can be drawn.
Figure 10. Stress in the Ni3Sn4 joint layer for four
different temperature on the hot side. Only limited
variations in stress level is observed.
Acknowledgements
The authors would like to acknowledge TECHNI AS,
TEGma AS and the Norwegian Research Council for
supporting this project (project No.: 244915).
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Figure 11. Stress in the Ni3Sn4 joint layer for varying
bond layer thicknesses. Insignificant variations in the
stress level is observed.
ISBN 978-0-9568086-1-5
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