Influence of Patterned Cu Roughness on
Cu-EMC Interface Adhesion Properties
Kaipeng Hu
MT 11.44
Internship report
Supervisors:
dr.ir. Olaf van der Sluis (TU/e, Philips Research)
ir. Sander Noijen (Philips Research)
Eindhoven University of Technology
Department of Mechanical Engineering
Division of Computational and Experimental Mechanics
Section Mechanics of Materials
November , 2011, Eindhoven
Contents
Abstract
i
1 Introduction
1
2 Theoretical Background
3
2.1
2.2
2.3
Cu-EMC Interface . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
Interface Fracture Mechanics . . . . . . . . . . . . . . . . . . . . . . . . . . .
Crack Kinking Theory . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
3
3
5
2.4
Adhesion Measurement Techniques . . . . . . . . . . . . . . . . . . . . . . . .
6
3 Experimental Analysis
8
3.1
Sample Preparation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
8
3.2
Four-point Bending Test . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
10
3.3
Failure Surface Investigation . . . . . . . . . . . . . . . . . . . . . . . . . . . .
11
3.4
Cross-section Investigation . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
12
3.5
Numerical Model . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
13
4 Results and Analysis
15
4.1
Load-displacement Curve . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
15
4.2
4.3
Numerical Simulation Results . . . . . . . . . . . . . . . . . . . . . . . . . . .
Surface Morphology . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
18
20
4.4
Failure Surface SEM Images . . . . . . . . . . . . . . . . . . . . . . . . . . . .
23
4.5
Cross-section SEM Images . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
24
5 Conclusions and Recommendations
5.1 Conclusions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
5.2 Recommendations . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
27
27
28
Bibliography
29
Acknowledgements
31
i
Abstract
Debonding of polymer-metal interfaces is one of the main failure modes observed in microelectronic components, which implies the adhesion at the interface is not sufficiently high.
Surface roughening is one generally accepted way to enhance interface adhesion. One of the
key mechanisms, besides the obvious increase in surface area and mechanical interlocking, is
the transition from adhesive to cohesive failure, i.e. crack kinking. In this report an adhesion
promotion method for Copper-Epoxy Moulding Compound (Cu-EMC) interfaces via simplified 2D microscopic groove-patterned roughness on the copper substrate is demonstrated.
The focus is to obtain reproducible crack kinking, assess the relevant failure mechanisms and
investigate the influence of surface roughness on Cu-EMC interface adhesion properties. The
specimens used in the experimental study are made of Cu-EMC bimaterial strips. Special
patterns composed of oriented parallel grooves of varying widths are etched on the copper
substrate surface. The specimens are debonded by means of the four-point bending test. At
first, an estimate of the interface fracture toughness is extracted from the load-displacement
curves using an analytical solution. Subsequently, simulations with cohesive zone elements at
the interface are performed to evaluate the interface fracture toughness with the consideration of large strain, plastic deformation and friction. Several unusual features of experimental
load-displacement curves are explained with the aid of simulation results. The failure surface and cross-section of copper substrate are investigated under both optical microscope
and scanning electron microscope (SEM). A detailed morphology of the copper leadframe is
deduced and the kinking fraction is calculated.
Chapter 1
Introduction
With continuous reduction of circuit sizes, the reliability and durability of polymer-metal
interface is a significant concern in microelectronic industry. It has been found that delamination of two material layers is a key failure mechanism of the polymer-metal interface [1].
Thus, during the past twenty years, considerable attention has been paid aiming at understanding of this failure mode. This includes analytical [2], experimental [3–5], and numerical
studies [6–8].
The occurrence of delamination depends on debonding driving forces, such as intrinsic residual
stress, thermal stress and external loading, and interface fracture toughness which can be
characterized by critical energy release rate (CERR) Gc . Therefore, reducing debonding
driving forces and increasing the interface fracture toughness are two strategies to improve the
interface reliability. This work focuses on the latter manipulation and elaborates an interface
fracture toughness promotion method by utilization of patterned surface roughness. Although
this a generally accepted method to enhance adhesion [9], quantitative understanding is still
lacking.
In order to improve the interface fracture toughness, the mechanisms of adhesion and cohesion
should be understood. The following interactions at various length scales play a role in
adhesion and cohesion of materials and interfaces [7].
1. Chemical interactions: Chemical interactions inside materials and between material
surfaces refer to primary bonds, i.e. covalent, ionic or metallic bonds. The interaction
scale is about 0.2 − 1.0 nm.
2. Physical interactions: Physical interactions refer to secondary bonds inside materials
and between interface, like the Coulomb force or Van der Waals force. Although in
general the magnitude of the physical interactions is weaker than the one of the primary
bond, it is formed at most interfaces while chemical interactions are not met. The
interaction scale is about 5.0 − 10.0 nm.
3. Mechanical interlocking: Mechanical interlocking is the interactions between material
surfaces or inside materials due to geometric effects at microscopic scale. The interlocking is considered to be the most significant mechanism in interface bonding. The scale
1
of surface roughness of metal-polymer interfaces, which is a typical interlocking feature,
is in the order of micrometers.
When a crack impinges on the polymer-metal interface, it may either propagate along the
interface or deflect into the polymer. Thus, debonding of polymer-metal interface might
involve both adhesive and cohesive failure [8]. The cohesive failure refers to the situation in
which the crack penetrates the polymer, which can be also called crack kinking [2]. Since the
fracture toughness of the polymer is usually much greater than the polymer-metal interface
fracture toughness, it is desired to obtain more cohesive failure near the interface to promote
the interface fracture toughness. Kim et al. [10] employed micro-patterned surface topography
at polymer-metal surface and found that the roughness triggered mechanical interlocking
provides an energy dissipating process whereby viscoelastic and plastic energy dissipation
and corresponding cohesive failure are required for crack propagation. Therefore, due to the
mechanical interlocking mechanism, increasing contact surface and crack kinking, roughened
surfaces are used commonly in practice to obtain better interface fracture toughness.
In this report, a method of enhancing interface fracture toughness in a Cu-EMC bimaterial strip through the introduction of well-defined roughness profiles at the copper leadframe
surface is introduced. Various patterns composed of oriented parallel grooves are designed
and etched on the copper leadframe surface prior to the EMC molding, aiming to form mechanical interlocking, increase contact area and trigger crack kinking. The focus is to obtain
reproducible crack kinking, assess the relevant failure mechanisms and investigate the influence of surface roughness on Cu-EMC interface adhesion properties. Firstly, some theoretical
background knowledge such as interface fracture mechanics, interface description and Gc
measurement techniques will be introduced in the forthcoming chapter. In Chapter 3 experimental procedures such as sample preparation, four-point bending test (4PB) and failure
surface investigation via both optical microscope and scanning electron microscope (SEM)
will be demonstrated in detail. Additionally, a numerical model is introduced for a better
understanding of the experimental results. The results and analysis for load-displacement
curves, surface morphology, optical microscope and SEM images and numerical simulations
will be presented in Chapter 4. Finally, concluding remarks and recommendations will be
given in Chapter 5.
2
Chapter 2
Theoretical Background
2.1
Cu-EMC Interface
Due to its low cost, ease of processing, and excellent electrical properties, EMC becomes
one of the most important materials used in encapsulation of the “System in Package” (SiP).
However, the relatively high thermal expansion of epoxy induces thermal stress in the package
during thermal cycling. Thus, inorganic fillers such as fused silica are used to lower the
thermal expansion of EMC. Regarding the leadframes, copper-based alloys are currently the
most widely used leadframe materials in microelectronic industry. It is ascribed to their
high thermal and electrical conductivities, and relatively low cost. Thus, the investigation of
Cu-EMC surface has received a significant attention [11–14]. H. Y. Lee et al. [11] reported
that chemical oxidation treatment of the copper-based leadframe before EMC molding could
potentially increase the interface fracture toughness. However, another interface fracture
toughness promotion method, i.e. surface roughening will be discussed here.
2.2
Interface Fracture Mechanics
To better understand the cracking behavior at Cu-EMC interface, basic knowledge of traditional interface fracture mechanics is required. Within the context of linear elastic fracture
mechanics (LEFM), it is assumed that the considered materials are all linear elastic and
isotropic and the fracture process zone is assumed to be small compared to all other dimensions. The foundation for linear elastic interface fracture mechanics is based on an asymptotic
analysis of the stress and strain fields near the tip of a crack [9]. The problem of interest
is illustrated in Figure 2.1. A semi-infinite crack with a straight front propagates along the
interface of two dissimilar linear elastic solids. The materials have the Young’s modulus Ei ,
the shear modulus µi and Poisson ratio νi , where i = 1, 2. The subscripts 1 and 2 represent
material 1 and material 2, respectively.
3
σ 22
x1
σ 12
σ 11
r
μ 1,ν 1
σ 11
σ 12
θ
σ 22
x2
μ 2,ν 2
Figure 2.1: Stress state at the interfacial crack tip [9]
The material is subjected to static remote loading, and is assumed to deform in plane strain.
Dundurs [15] introduced two elastic mismatch parameters
α=
E1 − E2
E1 + E2
β=
(1 − 2ν2 )/µ2 − (1 − 2ν1 )/µ1
,
2(1 − ν2 )/µ2 − 2(1 − ν1 )/µ1
(2.1)
where E i = Ei /(1 − νi 2 ). Evidently α is a measure of the relative stiffness of the two
materials while β is a rough measure of the relative compressibility of the two materials [15].
The condition α = β = 0 refers to a homogeneous condition.
For each material pair, a singular crack tip field exists at the crack tip according to linear
elasticity theory. For problems considered here, the normal and shear stresses of the singular
field acting on the interface a distance r ahead of the tip can be written in a compact form
σ22 + iσ12 =
(K1 + iK2 )riε
√
,
2πr
(2.2)
√
where i ≡ −1 and the crack tip singularity parameter ε (also called mismatch parameter)
depends on β according to
1
ε=
log
2π
1−β
.
1+β
(2.3)
For most of the bimaterials, the ε value is small and can be ignored. This leads to the fact that
the amplitude factors K1 and K2 can be interpreted as Mode I and Mode II stress intensity
factors, respectively. They are defined to be consistent with corresponding stress intensity
factors, KI and KII , which are for cracks in homogeneous problems. It implies that they also
depend linearly on the applied loads and the geometry. Thus, the energy release rate of the
crack in the interface is related to the interface stress intensity factor by
1
Gi =
2 cosh2 πε
1
1
+
E1 E2
(K12 + K22 )
(2.4)
Another important parameter representing the relative proportion of Mode I and Mode II
loading is the mode angle Ψ. For discussion purpose (as well as in many actual implementation
cases), it is assumed that β = 0. Thus
4
Ψ = tan
−1
σ12
σ22
(2.5)
When Ψ = 0◦ , it corresponds to pure mode I loading and Ψ = 90◦ corresponds to pure mode
II loading.
2.3
Crack Kinking Theory
As mentioned before, debonding of bimaterials consists of adhesive and cohesive failures.
He and Hutchinson [2] established an energy criterion to investigate the interface fracture
toughness promotion due to the surface roughness. From it, the conditions of the cohesive
failure, i.e. crack kinking, can be identified.
Given the fracture toughness of the interface Gic and that of the EMC Gec . The crack will
propagate along the interface if the condition Gi ≥ Gic is met. On the other hand, the crack
will deflect into the EMC if the energy release rate for the crack tip in the EMC, Ge , is
greater than Gec . Analogously, the crack will kink into the copper if Gc ≥ Gcc , where Gcc and
Gc are fracture toughness of copper and energy release rate for the crack tip in the copper,
respectively. Since the fracture toughness of the copper Gcc is much higher than those of the
interface and the EMC, the possibility of deflection into the copper is relatively small and it
will not be considered here. In the case where an interface crack kinks out of the interface
and into EMC the energy release rate Ge of the kinked crack may be determined from
Ge =
1 − ν1 2
2
(KI + KII
)
2µ1
(2.6)
Note that KI and KII are stress intensity factors of the kinking crack and are different from
the interface stress intensity factors K1 and K2 .
The energy release rate ratio GR is introduced in order to formulate the above discussion into
a mathematical form. The definition is as follows
GR =
Gi
Ge
(2.7)
Then the crack propagation path can be related to the energy release rate ratio GR in the
following way
Gi
Gic
>
⇒ crack propagates along the interface (interface failure)
Ge
Gec
Gi
Gic
GR =
<
⇒ crack deflects into the EMC (cohesive failure in EMC)
Ge
Gec
GR =
5
(2.8)
2.4
Adhesion Measurement Techniques
The CERR is dependent on the mode angle. Figure 2.2 present several basic configurations
for adhesion measurement under different mode angles. The double cantilever beam (DCB)
(a) and end notch flexure (ENF) (b) tests can induce pure Mode I and mode II delamination,
respectively. The mode angle Ψ for 4PB test (d) is around 45◦ [4] but is dependent on the
bulk materials and sample thickness. However, with the MMB test (c), a full range of mode
angles can be obtained by alteration of the distance d between two external loads. In this
project, configuration 4PB (d) test will be applied.
d
( a)
( b)
p
p
( c)
( d)
Figure 2.2: Schematic of different loading angles for interface delamination:(a) DCB test;
(b) ENF test; (c) mixed mode bending test; (d) 4PB test.
sudden drop
plateau
P/2
Pre-crack
b
h1
h2
h
1
2
initial slope
Pc
Load P
P/2
Crack
L
Knife Edge
Displacement δ
L
(a) Schematic illustration
(b) Typical load-displacement curve
Figure 2.3: Schematic illustration of the 4PB setup (a) and the typical load-displacement
curve (b).
In Figure 2.3(a) the schematic illustration of 4PB test is presented. The external load is P
and the dimensions of the sample are also indicated. The sample is a layered bimaterial strip
with different materials 1 and 2. Initially, it contains a pre-crack. When the load reaches
a critical value, the pre-crack will propagate perpendicularly towards the interface and grow
longitudinally along the weakest interface or deflect into the bulk materials. The load is
applied to allow crack propagation to occur, and stops when the crack tip approaches the
inner knife edges.
Figure 2.3(b) shows a typical load-displacement curve for a 4PB test. The initial slope of this
curve represents the bending stiffness of the specimen. When the load reaches a certain value,
6
the crack starts to propagate, which means a release of stress and energy. It can be related
to the sudden drop of the curve. After the drop, the curve has the trend of leveling off. If no
bulk deformation is considered, the energy given by the external load is all used for the crack
propagation, which explains the steady feature of the curve.
Indicated in the interface mechanics [4], when the crack length significantly exceeds the thickness of the EMC, the crack growth displays a steady-state feature. The constant load Pc ,
which is independent of crack length can be related to the CERR Gc [4]. Note that small
strain elasticity, plane strain conditions and symmetric delamination are assumed. Moreover, residual stress, plastic deformation and frictional dissipation are neglected during the
deduction of this equation,
3(1 − ν22 )Pc2 L2
Gc =
2E2 b2 h3
"
λ
1
− 3
1 η2
3
3
η2
η1 + λη2 + 12λ η1η+λη
2
#
,
(2.9)
where subscripts 1 and 2 denote material 1 and material 2, respectively. L is the distance
between inner and outer dowel pins (see Figure 2.3(a)); Pc is the load during steady propagation which can be obtained from the load-displacement (see Figure 2.3(b)); b and h are the
width and thickness of the sample. The non-dimensional parameters are:
λ=
E2 /(1 − ν22 )
,
E1 /(1 − ν12 )
ηi =
7
hi
(i = 1, 2)
h
(2.10)
Chapter 3
Experimental Analysis
3.1
Sample Preparation
The Cu-EMC sample used in this project is shown in Figure 3.1. It is a layered bimaterial
strip. The top layer with the notation of “1” is EMC. It has a thickness of h1 = 0.5 mm or
h1 = 1.0 mm. Two values are tried in order to trigger more delamination. While the bottom
layer, i.e. layer 2 is copper substrate with thickness of h2 = 0.2 mm. The length and width
are a = 48 mm and b = 8 mm, respectively.
In the partial enlarged view, the cross-section of the sample along y-direction is indicated.
The grooves are created in the x-direction. It is assumed that there is no geometry variation
in x-direction of the sample. Alternatively, all the cross-sections along the y-direction have
the same profiles. It is worth noting here that one reason for choosing this simplified 2D
geometry is the consideration of quantitative validation of numerical models to prevent any
ambiguities. In addition, the well-defined profiles make it easier for identification of failure
mechanisms compared to the stochastically distributed 3D surface roughness.
a
b
h1
h2
h
1 EMC
2 Copper
Pre-crack
z
y
x
Etched Copper Surface
Figure 3.1: Schematic of the Cu-EMC sample with pre-crack for 4PB test.
8
There are four main steps to prepare the samples: mask design, etching, molding and laser
cutting. The flow chart for these procedures is indicated in Figure 3.2.
A-view
W g1
W p1
W g1
B-view
W g2
A
W p2
W g2
B
(a) Mask design
(b) Etched copper surface (top view)
(c) Copper leadframe with molded EMC
(d) Laser cutting
Figure 3.2: Flow chart of the sample preparation. After the etching mask in (a) is designed,
two different ones are placed on the copper leadframe with size of 35 × 50 × 0.2 mm3 during
the etching process. The top view SEM image of copper surface after etching is shown in (b),
the grey lines are the grooves etched during the process. The EMC is molded subsequently
onto the copper substrate. (c) shows the badge of sample after molding. Finally, laser cutting
(d) is employed to obtain samples for 4PB test.
Firstly, etching masks for protection are designed. Since different surface profiles are fabricated, different masks are designed as well. Note that there are two profiles in one etching
mask. One example is shown in Figure 3.2(a). The profile of left and right part of the mask
is different, which can be further confirmed by the partial cross-section views A and B. The
variation is realized through the groove width Wg and plateau width Wp . After the etching
masks are designed, they can be reused.
Etching is the process of using strong acid or mordant to corrode the unprotected parts of
a metal surface to create a desired profile. In this project, the special pattern of roughness
at the copper surface is “grooves” with Wg ranging from 20 µm to 60 µm while the plateau
9
Figure 3.3: 4PB test setup
ranges also from 20 µm to 60 µm (see Figure 3.2(b)). Note that this picture is from the Set-1
samples. The grooves are not semi-circular. But samples in Set-2 and Set-3 are better. The
copper leadframe has a size of around 35 × 50 × 0.2 mm3 and two different etching masks are
placed on it in order to get four different profiles.
After the etching process, a layer of commercial black EMC with a thickness of 0.5 mm
or 1.0 mm is molded on the copper leadframe. It is necessary to clean the copper surface
with a sulphuric acid dip and plasma before molding in order to improve the conditions for
adhesion. The molding process lasts 180 s with the temperature of 180 ◦ C and the pressure of
200 bar. The post mold cure takes 4 hours at 175◦ C. Thus, badges with typical dimensions
of 35 × 50 × 0.7 mm3 (when h1 = 0.5 mm) or 35 × 50 × 1.2 mm3 (when h1 = 1.0 mm) are
obtained. The picture of the badge is shown in Figure 3.2(c).
Subsequently, four strips are laser cut from one badge. Finally, a pre-crack as shown in
Figure 3.1 is made by laser cutting.
The following table shows the material properties of two materials.
Table 3.1: Material properties of the copper leadframe and EMC.
Material 1 (EMC)
Material 2 (Cu)
3.2
Young’s Modulus E (MPa)
21240
123000
Poisson’s Ratio ν (-)
0.25
0.33
Four-point Bending Test
The interface fracture toughness, which is the critical energy release rate when fracture occurs,
is measured by four-point bending test. A picture of the setup is presented in Figure 3.3. The
samples are tested with a constant loading rate of 0.5 mm/min. The loads are measured with
a resolution of 0.01 N and digitally recorded to generate the load-displacement curve. The
test will be stopped when the crack approaches the inner knife edges (see Figure 2.3(a)). All
tests are conducted at room temperature.
10
During the experiment, three sets of samples, namely Set-1, Set-2 and Set-3, with different
surface morphologies are tested. Additionally, the thickness of EMC in Set-1 and Set-2 is
0.5 mm while in Set-3 it is 1.0 mm. Two samples in set-3 are glued with an additional
copper layer. The alterations are intended to increase the stored energy in the bulk materials
and thus trigger more delamination, which can eventually increase the possibility of crack
kinking. Except for two samples from Set-3, namely sample S1 and L1, all the samples
exhibit a delaminating interface. Thus, sample S1 and L1 are left out of the analysis. As will
be indicated in Section 4.3 samples in Set-1 show a small amount of crack kinking while Set-2
samples barely have crack kinking and Set-3 samples display a large area of it. The reasons
for this difference will be explained later.
3.3
Failure Surface Investigation
Since the 4PB tests are terminated when the cracks approach the inner knife edges, the
EMC layers are still partially attached to the copper leadframe. After the 4PB tests, the
remaining EMC layers are removed from the copper leadframe manually for the failure surface
investigation. The surface profiles of all the samples are investigated under optical microscope
and SEM. It is worth mentioning that the analyzed surface is indeed induced by the 4PB test
instead of the manual peeling.
residual EMC
Wg
wp
Copper
Figure 3.4: Schematic of the groove width Wg and plateau width Wp from copper leadframe
cross-sectional view. The residual EMC in the groove is indicated as well.
11
(a) Residual EMC on copper side
(b) Algorithm of kinking fraction
f computation
Figure 3.5: (a): Some residual EMC on the copper surface highlighted by white ellipse; (b):
illustration of the algorithm of computation of kinking fraction.
Firstly, the geometry of the copper surface is analyzed under optical microscope after the
4PB test. It consists of the measurement of the groove width Wg and plateau width Wp as
shown in Figure 3.4. For both measurements, 15 values from different locations are collected
while also the average and standard deviations are computed.
Secondly, the kinking fraction is estimated manually. Due to the etched grooves on the copper
surface, some EMC is left on the copper side (see Figure 3.5). Therefore on the EMC side the
strips are discontinuous, i.e. some portions are left on copper side. The total length Lt of the
strips on the EMC surface is determined. Then the length of the failure parts Lf is counted
(represented by red line in Figure 3.5). The kinking fraction f is calculated by f = Lf /Lt .
Note that images shown here are from the sample 2 in Set-1, which shows a small amount of
crack kinking. Samples in Set-2 and Set-3 show different results: Set-2 samples barely have
crack kinking while the grooves of Set-3 samples are fully filled with residual EMC.
Next the data obtained in the previous two steps is assembled and a correlation between the
surface morphology and the kinking fraction f is given.
Finally, the SEM is utilized to examine the failure surface of copper side aiming to have an
insightful investigation of the failure mechanisms.
3.4
Cross-section Investigation
Several cross-section samples are made and investigated with SEM in order to get more details
of the residual EMC in the etched grooves.
12
(a) Cutting
(b) Cross-section sample
Figure 3.6: (a) Cutting trajectory and view direction of the cross-section samples; (b) Crosssection sample of copper leadframe containing residual EMC.
The cut of the copper leadframe is chosen along the line where the density of the crack kinking
is relatively larger. The rectangular in Figure 3.6(a) shows the cutting trajectory. Attention
should be paid that the rectangular should be limited to the vicinity of the pre-crack to ensure
that the residual EMC is induced by the 4PB. The arrow indicates the view direction when
the cross-section sample is completed.
One of the cross-section samples is presented in Figure 3.6(b). A transparent polymer is
chosen as the embedding resin in order to obtain a clear view of the copper and EMC in
SEM.
3.5
Numerical Model
In this section numerical simulations will be performed aiming to obtain the interface fracture
toughness of several samples in Set-3. The analytical solution in (2.9) of the interface fracture
toughness is deduced under the assumptions of small strain, symmetric delamination, plane
strain conditions and elasticity [4]. However, in the real tests, these assumptions do not
always hold. Moreover, the residual stress in the bulk materials, plastic deformation of the
copper leadframe layer and frictional dissipation energy are not considered. These lead to the
fact that the experimental load-displacement curves (presented in Section 4.1) are different
compared to the typical load-displacement curve (see Figure 2.3(b)). Therefore, the analytical
solution from (2.9) is not reliable anymore. Besides a more accurate estimation of the interface
fracture toughness, it is also necessary to use a numerical model to analyze which factors
influence the load-displacement curves.
In the numerical model employed here, it is assumed that symmetric delamination, plane
13
Support
Initial delamination
Copper leadframe
EMC
Symmetry condition
Figure 3.7: Finite element model of the 4PB Cu-EMC sample.
strain conditions, large strain and plastic deformation exist and residual stress is not taken
into account. Only the effects of friction are investigated. The analytical solution from (2.9)
can be used as initial guess for the interface fracture toughness of the numerical simulations.
The constructed finite element model is indicated in Figure 3.7. With the assumption of symmetric boundary conditions, it only presents half of the sample. The supports are explicitly
modeled by rigid contact bodies. The mesh of the copper leadframe and EMC consists of 3000
and 5000 4-node quadrilateral plane strain elements, respectively. The interface fracture behavior is modeled by cohesive zone elements. Note that there is no initial crack in the model,
which implies that the force drop due to the onset of crack propagation will not be present
on the numerical load-displacement curves. Since our interest is on the crack propagation
regime, the absence of the initial crack is acceptable. However, there is a initial delamination
in the horizontal direction sa shown in the figure.
Simulations with various interface fracture toughness ranging from 60 J/m2 to 110 J/m2
and different friction coefficients are performed in commercial finite element software MarcMentat. These values are chosen due to the fact that the analytical estimation is 90 J/m2
(shown in Section 4.1). Various values are used to fit the experimental load-displacement
curves aiming to obtain a more accurate solution for the interface fracture toughness.
14
Chapter 4
Results and Analysis
In this chapter, experimental and numerical results will be presented. Firstly, load-displacement
curves from 4PB tests are illustrated. It is shown that these curves have different shapes
compared to the typical load-displacement curve. They can be utilized to extract analytical
solutions of the interface fracture toughness according to (2.9) as initial guess for the simulations. Subsequently, results from numerical simulations with a consideration of nonlinearity
are shown aiming to understand the unusual features of the experimental load-displacement
curves. Then the surface morphology of the samples is analyzed in order to relate roughness
profile to crack kinking fraction f and interface fracture toughness. It turns out that two
categories of crack kinking are observed, namely micro-kinking and intended-kinking and the
intended-kinking is more favorable for the profiles with smaller grooves. Finally, SEM images
of failure surface and cross-section investigation are exhibited and analyzed. They reveal that
besides the obvious increase of contact area with a factor of approximately 1.5, mechanisms
like mechanical interlocking and crack kinking also contribute remarkably to the promotion
of the adhesion.
4.1
Load-displacement Curve
In Figures 4.1, 4.2 and 4.3 the load-displacement curves for three sets of samples are illustrated.
The experimental data for the three sets of samples are listed in Table 4.1, Table 4.2 and
Table 4.3. In the tables, h, B and Pmax represent sample thickness, initial slope and maximum
load, respectively. Obviously, the curves obtained in this project have different behaviors
compared to the typical load-displacement curve depicted in Figure 2.3. Instead, two regions
are distinguished:
1. initial slope regime due to the bending stiffness of the sample
2. delamination regime during which the load either decreases or increases
The first part is straightforward and is only influenced by mechanical properties of the bulk
materials, thickness of the layers and the geometry of the pre-crack. At this stage, the
15
samples deform elastically accumulating strain energy for subsequent cracking. The typical
stiffness for Set-1 and Set-2 is around 15.00 N/mm. For the Set-3 samples, the thicker EMC
and additional glued copper layer lead to higher sample stiffness ranging from 45 N/mm to
62 N/mm (see Table 4.3). The second regime, which refers to the crack propagation part,
displays a complicated behavior. Instead of forming a steady plateau as the classical solution
exhibits, some curves have an increasing trend at the beginning and decrease after reaching
peaks, the others keep increasing until the end of the test. These unusual features are caused
by different nonlinearity factors, which will be further discussed in the next section with the
aid of numerical simulations.
Note that at the end of the test, sample C2.1 slips off the set-up, one can see there is a
sharp drop of the curve in Figure 4.2. Regarding sample W2 and W4 in Figure 4.3, additional
copper layers of thickness 0.2 mm are glued, which clearly enhances the force levels.
Another apparent nature of all the curves is the serrated shape. Two stages on the curves
can be distinguished (see Figure 4.2). The scale of the serrated shape from the first phase is
smaller than the second one. The distance between the peaks in the first phase is typically 50
µm, while it is about 125 µm in the second phase. Moreover, the amplitude of the oscillation
in the second phase is larger than that of the first phase. It will be indicated later that
the width of the grooves etched on the copper surface is in the order of 50 µm. Therefore,
it is possible that the first phase of the serrated shape is caused by the complicated crack
propagation path induced by the roughened copper surface. Regarding the second phase, it
is assumed that the sliding of the samples is responsible for it. In addition, it is found that
this could also be attributed to unstable operation of the loading motor.
Solution from (2.9) and (2.10) is used as an initial guess for the interface fracture toughness
of the numerical simulations. For all the samples except for W2 and W4, the constant load
(Pc ) is estimated as 6.5 N. Then the CERR of the sample is around 90 J/m2 . Compared to
the CERR of the unetched samples, which is around 10 J/m2 , evaluated by M. Kolluri [16],
there is a pronounced improvement of the toughness with a factor of nine. Recall that the
contact area increase with a factor of about 1.5, there is additional dissipative mechanisms,
which will be demonstrated later.
16
9
8
Table 4.1: Data of Set-1 samples.
Load, P(N)
7
6
5
4
Sample 1
Sample 2
Sample 3
Sample 4
Sample A
Sample B
Sample D
3
2
1
0
0
1
2
3
4
5
6
Displacement, δ(mm)
Figure 4.1:
Sample
h (mm)
B (N/mm)
Pmax (N)
A
B
D
1
2
3
4
0.7
0.7
0.7
0.7
0.7
0.7
0.7
13.78
15.22
15.44
15.49
15.39
14.96
14.89
6.58
6.88
7.33
7.46
7.67
8.42
7.87
Load-displacement curves for sample Set-1.
8
Table 4.2: Data for Set-2 samples.
7
Load, P(N)
6
2nd stage
1st stage
5
4
3
Sample C1.1
Sample C1.2
Sample C1.3
Sample C1.4
Sample C2.1
Sample C2.2
Sample C2.3
Sample C2.4
2
1
0
0
1
2
3
4
5
6
Displacement, δ(mm)
Figure 4.2:
Sample
h (mm)
B (N/mm)
Pmax (N)
C1.1
C1.2
C1.3
C1.4
C2.1
C2.2
C2.3
C2.4
0.7
0.7
0.7
0.7
0.7
0.7
0.7
0.7
16.35
17.40
16.74
17.49
14.75
14.79
17.20
17.21
6.95
7.82
7.68
7.30
6.67
7.05
7.45
6.80
Load-displacement curves for sample Set-2.
40
35
Load, P(N)
30
Table 4.3: Data for Set-3 samples.
25
Sample S4
Sample L2
Sample L3
Sample W2
Sample W4
20
15
10
5
0
0
0.5
1
1.5
2
2.5
3
3.5
4
Displacement, δ(mm)
Figure 4.3:
Load-displacement curves for sample Set-3.
17
Sample
h (mm)
B (N/mm)
Pmax (N)
S4
W2
W4
L2
L3
1.2
1.4
1.4
1.2
1.2
45.24
56.75
62.86
47.69
48.20
7.65
37.01
34.05
8.24
7.38
Figure 4.4: Tested samples with permanent deformation.
4.2
Numerical Simulation Results
In order to identify the different phenomena on load-displacement curves mentioned in previous section and obtain a more accurate estimation of the interface fracture toughness, numerical simulations are performed. Load-displacement curves from numerical simulations are
presented in Figure 4.5. Results of various interface fracture toughness with friction coefficient
η = 0, η = 0.25 and η = 0.50 1 are shown in Figure 4.5(a), Figure 4.5(b) and Figure 4.5(c),
respectively. In Figure 4.5(d), interface fracture toughness values and different friction coefficients are indicated. In all figures, an additional horizontal line with Gc = 80 J/m2 and load
P = 6.18 N derived from (2.9) is added to indicate the deviation of the numerical simulation
and the analytical solution. It can be observed that as the η value increases, the deviation
also increases. Two phases of the curves can be recognized, namely the initial slope and
the crack propagation regime. The oscillations in the crack propagation regime are possibly
induced by the element size in the mesh.
Figure 4.5(a) reveals that when friction is not considered, the crack propagation regime displays a plateau. With an increasing interface fracture toughness, the height of the plateau
on the load-displacement curves increases. The gray dash line (Gc = 100 J/m2 , h = 1.2
mm) along with the gray dot line (Gc = 100 J/m2 , h = 0.7 mm) show that the thickness
only alters the initial slope of the curves, the crack propagation regime is not influenced
when the friction is neglected. However, when the friction coefficient is changed to η = 0.25,
Figure 4.5(b) indicates that the thickness of the sample has influence on the crack propagation region. It is further shown that the friction changes the shape of the curves. The lines
with high interface fracture toughness attain an increasing-decreasing shape, which is also
observed in the experimental load-displacement curves. As to the Gc = 60 J/m2 , it has an
increasing-decreasing-increasing shape, meaning that interface with lower toughness is more
sensitive to the friction, which is further confirmed by Figure 4.5(c). Figure 4.5(d) illustrates
1
It is the friction between the support and the bulk materials that is considered here.
18
the influence of the friction coefficients. An increasing friction coefficient increases the force
level and makes the curves steeper.
11
11
10
10
9
9
8
8
7
7
Load, P(N)
Load, P(N)
With these numerical results, the interface fracture toughness of three samples from set-3 are
evaluated. Sample L2, L3 and S4 correspond to Gc value of 65 J/m2 , 70 J/m2 and 80 J/m2 ,
respectively. It can be seen that the interface fracture toughness values estimated from the
numerical simulations are smaller than the analytical solution. This is attributed to the fact
that during the derivation of the analytical solution (2.9) none of the nonlinearity sources
such as large strain and plastic deformation are considered.
6
5
4
Gc=60,h=1.2
Gc=70,h=1.2
Gc=80,h=1.2
Gc=90,h=1.2
Gc=100,h=1.2
Gc=110,h=1.2
Gc=100,h=0.7
3
2
1
0
0
1
2
3
4
6
5
4
Gc=60,h=1.2
Gc=80,h=1.2
Gc=90,h=1.2
Gc=100,h=1.2
Gc=110,h=1.2
Gc=100,h=0.7
3
2
1
0
5
0
1
Displacement, δ(mm)
3
11
11
10
10
9
9
8
8
7
7
6
5
4
3
1
0
1
2
3
5
6
5
4
Gc=80,h=1.2
Gc=80,h=1.2
Gc=80,h=1.2
Gc=110,h=1.2
Gc=110,h=1.2
Gc=110,h=1.2
3
Gc=80,h=1.2
Gc=90,h=1.2
Gc=100,h=1.2
Gc=110,h=1.2
2
4
(b) friction coefficient η = 0.25
Load, P(N)
Load, P(N)
(a) friction coefficient η = 0
0
2
Displacement, δ(mm)
4
2
1
0
5
Displacement, δ(mm)
0
1
2
3
4
5
Displacement, δ(mm)
(c) friction coefficient η = 0.5
(d) combination
Figure 4.5: Load-displacement curves from numerical simulations and the solid line corresponds to the Gc = 80J/m2 and P = 6.18 derived from (2.9)
Note that symmetric delamination, plane strain conditions, large strain and plastic deformation (a hardening law is defined in the material model) are assumed and residual stress
is not taken into account in this model. Only the effects of friction are investigated. From
the numerical simulations, the complicated phenomena of the experimental load-displacement
curves can be explained by the interplay of the following factors:
19
1. Increasing portion of the Mode II fracture during the test. As mentioned earlier, normally the 4PB test consists equal normal and shear displacements on the interface.
Nevertheless, this hypothesis does not hold anymore when the crack approaches the
inner loading line. The Mode II portion will increase at the vicinity of the inner loading
line, which causes the increase of the mode angle Ψ [4]. Note that the interface fracture toughness is not a single material parameter, rather it is a function of the relative
amount of Mode II to Mode I acting on the interface (see Section 2.2). It has been
proven by Wang and Suo [17] that an increasing mode angle Ψ leads to the growth of
the interface fracture toughness, and thus a load increase. With an increasing Mode II
loading magnitude, the frictional contact between the fracture surfaces and plasticity
in adjacent layers also increases, which will also raise the load.
2. Plastic deformation. It can be seen from Figure 3.3 and Figure 4.4 that the samples
show permanent deformation after the test. It indicates that the work of separation is
not only used for the crack propagation, but also for the plastic bending of the sample.
However, numerical simulation reveals plastic deformation will actually decrease the
load [18]. This observation requires additional verification.
3. Sliding behavior of the samples from the knife edges. As the experiment goes on, the
deformation of the samples becomes large, which makes the samples slide from the knife
edges. From (2.9), with a certain interface fracture toughness Gc , the sliding behavior
of the samples will decrease the effective loading span L, thus increase the load needed
for a certain crack propagation.
4. Frictional contact. Numerical simulations show that the consideration of frictional dissipation energy will increase the force levels.
5. Residual stress and symmetry. Although these factors are not considered here, these
might affect the load-displacement curves.
4.3
Surface Morphology
The surface morphology investigation is carried out to relate the surface profiles to crack
kinking and interface fracture toughness. The results for the surface investigation are listed
in Table 4.4, Table 4.5 and Table 4.6. Since sample A and sample B in Set-1 are not etched
properly as shown in Figure 4.7, they are removed from the analysis. Furthermore, the error
bar plot for this data is presented in Figure 4.6.
Table 4.4: Surface profile of sample Set-1.
Sample ID
1
2
3
4
C
D
Wg (µm)
31.30±1.48
33.49±1.15
37.49±0.65
35.12±1.67
22.15±0.99
20.80±0.76
Wp (µm)
58.64±0.67
57.37±1.66
51.58±0.76
53.96±1.41
55.98±1.06
58.08±0.82
20
Kinking fraction [%]
12.86
14.82
3.88
5.41
1.31
2.80
Wg /Wp ratio
0.53
0.58
0.73
0.68
0.39
0.35
Table 4.5: Surface profiles of sample Set-2.
Sample ID
C1.1
C1.2
C1.3
C1.4
C2.1
C2.2
C2.3
C2.4
Wg (µm)
56.61±0.94
56.90±0.62
61.48±1.20
60.23±1.21
55.96±0.72
55.22±0.86
62.59±0.95
61.33±0.90
Wp (µm)
33.16±1.39
32.75±0.62
28.88±1.37
29.24±0.61
24.50±1.33
25.46±1.18
18.26±0.88
19.07±1.09
Kinking fraction [%]
0
0
0
0
0
0
2.54
1.24
Wg /Wp ratio
1.70
1.70
2.10
2.07
2.30
2.20
3.50
3.21
Table 4.6: Surface profile of sample Set-3.
Sample suffix
1
2
3
4
Wg (µm)
31.29±1.29
37.12±0.75
38.18±1.65
46.53±1.96
Wp (µm)
60.15±1.21
53.08±1.75
44.69±0.58
33.72±1.79
Kinking fraction
full
full
full
full
60
Wg /Wp ratio
0.52
0.70
0.85
1.37
65
60
55
55
Groove width
Plateau width
50
Width [µm]
Width [µm]
50
45
40
35
Groove width
Plateau width
45
40
35
30
30
25
25
20
20
0
1
2
3
4
5
6
15
7
Sample ID
0
1
2
3
4
5
6
7
8
9
Sample ID
(a) Error bar plot for sample Set-1
(b) Error bar plot for sample Set-2
Figure 4.6: Error bar plot for surface profiles of two sets of samples.
From the tables and figures given above, it can be seen that the crucial difference of sample
Set-1 and sample Set-2 is the values of the Groove/Plateau ratio. In Set-1 the width of the
plateau WP is 2-3 times as large as that of the groove WG and in Set-2 it is vice versa.
From the kinking fraction of Table 4.4 and Table 4.5, it can be concluded that kinking is more
favorable for the profiles with smaller grooves. Kim et al reported an opposite results [10].
Note that the relatively higher value of kinking fraction for sample 1 and sample 2 is due to
the occurrence of the edge effect (more kinking at the edge, see Figure 4.8). As to the sample
Set-3, the thickness of EMC is increased from 0.5 mm to 1.0 mm. Apparently, the thicker
EMC helps to form more kinking.
21
The key conclusion is that the simplified 2D profiles are able to trigger crack kinking. Due
to the limited variations, which combination of groove width and plateau width is optimal
for crack kinking still remains an unknown. Another interesting thing is that Set-2 samples
barely have crack kinking while the grooves of Set-3 samples are fully filled with residual EMC.
However, such a significant difference is not observed for the interface fracture toughness
values according to the load-displacement curves in Figure 4.2 and Figure 4.3. It is expected
that crack kinking is not the only mechanism that contributes to the enhancement of the
adhesion. It is necessary to analyze the failure surface and cross-section of Set-2 and Set-3
samples in SEM to obtain more details.
(a) Sample A
(b) Sample B
Figure 4.7: Optical microscope images of failure surface at copper side for sample A (a) and
sample B (b) .
(a) Sample 1
(b) Sample 2
Figure 4.8: Edge effect of sample 1 and sample 2.
22
4.4
Failure Surface SEM Images
In Figure 4.9 four SEM images of the copper leadframe failure surface from sample Set-2 and
Set-3 are shown. Figure 4.9(a) and Figure 4.9(b) are back scattered electron detector (BSED)
image and Everhart-Thornley detector (ETD) image, respectively, and both from the Set-2
sample C2.3. These images illustrate the typical topography of the failure surface, exhibiting
residual EMC on the copper side.
In the BSED image, two major colors, namely black and white, can be distinguished. The
black portion represents residual EMC, while the white counterpart stands for exposed copper
surface. For the residual EMC, it can be further divided into two categories: micro-kinking
and intended-kinking. The micro-kinking has a discontinuous feature, stochastically spreading
over the copper surface, both in the groove and on the plateau. The intended-kinking blocks,
which have a relatively larger dimension compared to the micro-kinking, are only located in
the grooves. Note that micro-kinking is believed to be induced by the small scale roughness
and the physical interactions between bulk materials. These imply that the secondary bonds
across the interface are strong enough to cause crack deflection into the epoxy resin, resulting
in a partial cohesive failure mode. The intended-kinking is the cohesive failure mode which
is present when a well-defined roughness pattern is applied on the copper surface.
For the ETD image, it is worth mentioning that there are only micro-kinking in the image
and the white spots in the field are induced by the electron charging. A clear difference
between the topography in the groove and on the plateau can be recognized. The density
of the micro-kinking in the groove is much larger than that on the plateau. Almost 95% of
area in the groove is covered by the residual EMC. The reason for this is a consequence of
the different asperity at these two areas. Since the grooves are produced by etching, they
are much rougher than the surface on the plateau (see Figure 4.10). It is not surprising that
more micro-kinking in the grooves is present. The higher number of EMC remnants in the
grooves also indicate the frictional resistance during the pull out of the interlocked EMC.
Note that Figure 4.9(b) is a representative image of the Set-2 sample failure surface, showing
micro-kinking in the whole view. However, Figure 4.9(a) displays a typical location where
intended-kinking and micro-kinking both exist.
Figures 4.9(c) and Figure 4.9(d) indicate the failure surface of sample S4 from Set-3. It can be
seen that the grooves on the copper substrate are almost fully filled with the residual EMC,
showing that numerous intended-kinking occurred during the 4PB test for this sample. The
right part of the Figure 4.9(c) is the place where the pre-crack is located. This part is included
in order to indicate that the image is taken in the vicinity of pre-crack, implying that the
failure surface in the view is induced by the 4PB test instead of the peeling after the test.
Figure 4.9(d) also displays a small portion of adhesive failure. In the groove, micro-kinking
can be found as well.
Recall that the contact area increases with a factor of about 1.5 and the interface fracture
toughness increases with a factor of nine. Now it can be concluded that the crack kinking is
a main contributor to the dissipative energy. The observed micro-kinking partially explains
similar load-displacement curves of Set-2 and Set-3 samples. The micro-kinking, one type of
cohesive failure mechanisms, caused by the molecular decohesion of the EMC, also consumes
large amount of energy for crack propagation. It should be emphasized that the dense micro23
kinking can play an equivalent role as intended-kinking in the adhesion promotion.
(a) BSED image of failure copper surface from (b) ETD image of failure copper surface from
Set-2 sample C2.3 .
Set-2 sample C2.3.
(c) BSED image of failure copper surface from (d) BSED image of failure copper surface from
Set-3 sample S4.
Set-3 sample S4 with higher manification.
Figure 4.9: SEM images of failure surface from sample Set-2 and Set-3.
4.5
Cross-section SEM Images
Figures 4.10(a) and (b) present the cross-section images of the copper leadframe of sample C2.3
in Set-2. As shown in the images, different amounts of EMC are left in the copper grooves.
Clearly, the cross-section of the periodic grooves is semicircular due to the isotropic etching
process. The gray lines represent the cohesive failure path. If the residual EMC is caused by
adhesive failure, it is not supposed to appear in the cross-section image, since it should be
gone together with the peeled-off EMC. Due to the cutting during the cross-section sample
24
preparation, the residual EMC in (b) is rotated and new cracks are formed which separate
the residual EMC from the bottom of the grooves. The circles in the images represent the
inorganic fillers aiming to lower the thermal expansion and improve the mechanical properties
of EMC. From the pictures it can be observed that the amount of the residual EMC is different
in different grooves. This is assumed to be a consequence of the distribution and different
geometry of the inorganic fillers. The difference in the microstructure of the residual EMC
might lead to the oscillations on the load-displacement curves. However, the general trend
of the curves maintains uninfluenced, implying that the interface fracture toughness is not
affected.
The cross-section image Figure 4.10(c) from sample S4 in Set-3 reveal that the EMC is
completely filled into the etched micro-grooves, implying that mechanical interlocking of the
EMC into the copper surface cavities was completely attained at the micro groove-patterned
Cu-EMC interface. Note that this image is taken at the area where the crack has not reached.
Figure 4.10(d) shows a detail of the crack propagation path where the crack propagates from
right to left.
25
(a) Cross-section 1 from sample C2.3 in Set-2 (b) Cross-section 2 from sample C2.3 in Set-2
(c) Cross-section image of undebonded part (d) Crack propagation path from sample S4
from sample S4 in Set-3.
in Set-3.
Figure 4.10: Cross-section SEM images of the samples from both Set-2 and Set-3.
26
Chapter 5
Conclusions and Recommendations
5.1
Conclusions
In this work, the interface fracture toughness of Cu-EMC bimaterial strip with well-defined
microscopic roughness pattern on copper surface is investigated experimentally and numerically. 4PB test is adopted in the experiment. The main conclusions are summarized below:
1. Three basic failure modes for 4PB experiment during the crack propagation process are
observed (see Figure 5.1). The first one is pure cohesive failure mode, which is called
intended-cracking in the preceding chapters. The second one and the third one are the
combinations of cohesive failure (only micro-kinking) and adhesive failure (clear delamination of bulk materials). However, the second one presents at the groove surface while
the third one occurs at the plateau surface. The results of failure surface SEM investigation presents a clear topology difference between the etched grooves and plateau.
The density of micro-kinking in the grooves is much larger due to the rougher groove
surface derived from the chemical etching.
2. The simplified 2D profiles are able to trigger reproducible crack kinking. Due to the
limited variations, which combination of groove width and plateau width is optimal
for more crack kinking still remains an unknown. However, the evaluation of kinking
fraction f shows that the profile of the copper leadframe surface has an influence on the
kinking fraction f . The intended-kinking is more favorable when the groove width Wg
is smaller than the plateau width Wp .
3. The load-displacement curves from 4PB show that the interface does not debond in a
steady manner. With the interplay of friction dissipation, plastic deformation, sliding
behavior, increasing portion of the Mode II fracture and residual stress, the curves
display an obvious discrepancy from the typical load-displacement curve.
4. The interface fracture toughness of the Cu-EMC sample is measured to be around 90
J/m2 , determined from analytical method. With the aid of simulations, the value is
finally determined between 65 J/m2 and 80 J/m2 . The difference indicates that the
analytical solution is an overestimation of the interface fracture toughness. It is six to
27
eight times as large as the counterpart of samples with unroughened copper leadframe.
The contact area increases only with a factor of around 1.5. Thus, it can be concluded
that the crack kinking is the main contributor to the enhancement of the interface
fracture toughness.
5. The numerical simulations show that the thickness of the EMC layer does not affect
the crack propagation region when the friction is not considered. As the friction increases, the curves have an increasing-decreasing trend, meaning that the force levels
are enhanced.
6. The cross-section images suggest that mechanical interlocking of the EMC into the
copper surface cavities was completely attained at the micro groove-patterned Cu-EMC
interface.
7. Experimental results show that the mechanisms of interface fracture toughness enhancement caused by metal surface topography modification in Cu-EMC bonded joints is not
only the obvious contact area and mechanical interlocking, but also the transition from
adhesive to cohesive failure. The cohesive failure can be further divided into microkinking and intended-kinking. It should be emphasized that the dense micro-kinking
can play an equivalent role as intended-kinking in the adhesion promotion.
1
3
2
Figure 5.1: Three failure modes during the crack propagation.
5.2
Recommendations
1. Instead of 4PB test, other adhesion test methods, such as mixed mode bending test [16],
can be adopted. Since the interface fracture toughness increase monotonically as the
mode angle Ψ increases [19]. By increase of the mode angle, more delamination can be
obtained. Alternatively, samples with thicker copper layer can be fabricated. Thus, the
influence of the thickness of bulk materials can also be further investigated.
2. Due to the small size of the groove-pattern, the fracture behavior at the interface is
not the same as the planar interface. In-situ experiments can be performed in order
to obtain more insights of the crack propagation process in the bimaterial patterned
interface and used as validation of the numerical models that are developed.
3. More tests are needed to correlate the surface topology and interface fracture toughness
Gc in order to obtain adhesion enhancement by controlling the surface profiles.
28
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Acknowledgements
I would like to express my gratitude to all those who helped to complete this internship
report. First, I want to offer my sincerest gratitude to one of my supervisors, Olaf van der
Sluis, for offering me the chance to do the internship in Phillips Research in Eindhoven.
I am grateful for his useful guidance, valuable encouragement and consistent patience. I
would like to thank Sander Noijen for offering me a lot of help in numerical parts. Many
thanks to Peter Timmermans from Philips Research, for supporting me in arrangement of
the experiments and helpful discussion during the weekly meetings. Furthermore, I would like
to thank Ron Hovenkamp for his help during the experiments in the Mechanical Test Lab in
Philips Research. Finally, I would like to thank Marc van Maris for his experimental guidance
about how to operate the advanced equipments in the Multi Scale Lab of the Technical
University of Eindhoven.
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