Indian Journal of Engineering & Materials Sciences
Vol. 17, June 2010, pp. 186-198
Damage prediction in glass/epoxy laminates subjected to impact loading
Ramazan Karakuzua, Emre Erbil
a & Mehmet Aktas
b*
aDepartment of Mechanical Engineering, Dokuz Eylül University, 35100, Izmir, Turkey b Department of Mechanical Engineering, Usak University, 64300, Usak, Turkey
Received 31 December 2009; accepted 31 May 2010
In this paper, the impact behaviour of glass/epoxy laminated composite plates with [0/±θ/90]S fiber orientation is
investigated numerically at equal energy (40 J), equal velocity (2 m/s) and equal impactor mass (5 kg). In order to examine
the stacking sequence effect, five different ±θ fiber directions are chosen as 15°, 30°, 45°, 60° and 75°. Three different plate
thicknesses as 2.9 mm, 5.8 mm, and 8.7 mm are also selected to survey the thickness effect on impact behaviour of
glass/epoxy composite plates. The overlapped delamination area is obtained for composite plates with different stacking
sequences and different thicknesses. A transient finite element code 3DIMPACT is used for numerical analyses. In this code
an eight-point brick element and the direct Gauss quadrature integration scheme are used through the element thickness to
account for the change in material properties from layer to layer. The Newmark scheme is also adopted to perform time
integration step by step. In addition, a contact law incorporated with the Newton-Raphson method is applied to calculate the
contact force during impact. Numerical results are compared with the experimental study and it has been seen that they are
in good agreement with the experimental results.
Keywords: Damage prediction, Low velocity impact, Glass/epoxy, Thickness effect, Stacking sequence effect,
Delamination
In recent years the use of composite materials has
become increasingly common in a wide range of
structural components, engineering applications,
aerospace, automotive, defense, and sports industries.
Composite materials have numerous advantages over
more conventional materials because of their superior
specific properties; such as high strength and stiffness
to weight ratio, improved corrosion, and
environmental resistance, design flexibility, improved
fatigue life, potential reduction of processing,
fabrication and life cycle cost.
Despite these advantages, laminated composites
can be susceptible to damages under transverse
impacts. The various damages, such as matrix cracks,
delaminations, fiber fracture, fiber-matrix debonding
and fiber pull-out can occur during impact event.
These damages cause considerable reduction in
structural stiffness, leading to growth of the damage
and final fracture. Therefore, the impact response of
fiber reinforced laminated composites has been an
important area of research for a long time. A number
of studies in this field have already been reported in
literature1-3
.
Abrate4 has used four mathematical impact models
such as spring mass, energy balance, complete models
and an impact on infinite plate model for the analysis
of the dynamic and quasi-static behaviours of
composite structures. Results showed that the spring
mass and energy balance models might be suitable in
the quasi-static case. It is also suggested the complete
model to take into account the full dynamic behaviour
of the plate. Schoeppner and Abrate5 have
investigated the delamination threshold load for low
velocity impact on graphite/epoxy, graphite/PEEK,
and graphite/BMI composite laminates for the number
of layers ranging from 9 to 96 plies. Result exhibited
that the graphite/BMI laminates have a higher
delamination threshold load and higher damage
resistance over the entire range of layer numbers. Hou
et al.6 have described an improved delamination
criterion for laminated composite structures and its
implementation to LS-DYNA3D. The influences of
high local interlaminar shear stress induced by matrix
cracking and fiber failure on delamination have also
been taken into account. Results showed that the
delamination near the impacted surface is induced by
contact stress due to the Hertzian effect.
Mikkor et al.7 have investigated the effect of
various parameters including the magnitude of the
preload, impact velocity and specimen geometry, to
predict the behaviour of preloaded carbon/epoxy _____________
*Corresponding author (E-mail: [email protected])
KARAKUZU et al.: DAMAGE PREDICTION IN GLASS/EPOXY LAMINATES
187
plates subjected to a range of low energies by using
explicit finite element model, which incorporates a bi-
phase material degradation model. The results showed
that at lower preloads; the damage size and residual
strength reduction increased with increasing impact
velocity. At higher preloads, increasing the impact
velocity did not appear to affect significantly either
the damage size or the residual strength prior to
catastrophic failure. Breen et al.8 have simulated the
dynamic and quasi-static impact on CFRP laminates
to obtain the residual tensile strength by using
ABAQUS/EXPLICIT software. The approximately
20% lower residual strength was obtained in
dynamically impacted specimens compared to the
simulated specimens using the quasi-static test.
Hosseinzadeh et al.9 have investigated the damage
behaviour of thin and thick glass/epoxy, carbon/epoxy
and sandwich structures including carbon/glass/epoxy
under different impact energies experimentally and by
using ANSYS/LS DYNA software. The thick
glass/epoxy plate has shown the best stability for all
impact energies. The carbon/epoxy plate has shown
good structural resistance compared to other
materials. The carbon/glass/epoxy plate was proposed
as optimized material in terms of strength and weight
reduction. Aslan et al.10-12
have conducted a few
numerical and experimental studies to investigate the
stacking sequences, dimensions and thicknesses of
composite plates, impactor mass and velocity effects
on the impact behaviour of glass/epoxy composite
laminates. Results showed that the impact forces,
deflection and delamination area were proportional to
the impactor velocity and mass. In addition, they have
discovered that the smaller width of the rectangular
composite laminate has the higher contact duration.
Zhang et al.13
have developed a numerical model to
predict the damage initiation and propagation for
cross-ply carbon/epoxy composite plates subjected to
low velocity impact. The model was implemented
into ABAQUS/Explicit commercial finite element
package. It can be said from the numerical study that
the numerical model can reliably predict the
characteristics of the low velocity impact damage of
composite structures. Li et al.14
have developed a
numerical model to simulate the damage process of
cross-ply carbon/epoxy composite laminates under
low velocity impact. In this model, the 9-node
Lagrangian element of the Mindlin plate was
employed to consider the large deformation. Many
critical aspects have been verified through many
previous experimental and numerical results. Her and
Liang15
have investigated the effect of shell curvature,
boundary conditions and impactor velocity on the
graphite/epoxy composite plates subjected to low
velocity impact by using ANSYS/LS-DYNA
software. Results showed that the contact force is
proportional to the impactor velocity. However, the
contact duration is dependent on the stiffness of the
laminated structure, such as the curvature and
boundary conditions. Moura et al.16,17
have proposed a
numerical model to predict the damage mechanisms
in low velocity impact of cross-ply carbon/epoxy
laminate. The predicted results were compared with
the experimental ones. Good agreement between both
analyses was obtained for shape, orientation and size
of the delamination.
Johnson et al.18
have presented a material failure
model, which includes both intraply damage and
delamination, for carbon/epoxy composites plates.
This model was implemented into the dynamic finite
element code PAM/CRASH. The code was applied at
different velocities by a steel impactor. A comparison
of structural response and failure modes from
numerical simulation and impact tests was given at
low impact energies. Naik et al.19
have investigated
the impact behaviour of woven glass/epoxy and
T300/5208 carbon/epoxy laminated composite plates
under low-velocity impact by using a modified
Hertzian contact law. The results showed that the
maximum displacement and maximum in-plane
failure decrease by increasing the in-plane modulus of
elasticity.
Davies et al.20
have predicted the threshold impact
energy for the onset of delamination in quasi-isotropic
carbon/epoxy laminates under low-velocity impact.
The quasi-static prediction was compared with the
corresponding experimental results for various
laminates with different sizes and different boundary
conditions. It was found that the theoretical prediction
is in good agreement with the experimental data.
Tiberkak et al.21
have used Mindlin’s plate theory,
along with a modified Hertzian contact law to
describe the impact behaviour of carbon/epoxy
laminated composite plates subjected to a central
impact in a spherical projectile. The effects of some
parameters of the impactor and the composite plate,
such as mass and velocity of the impactor, as well as
the stacking sequence, boundary conditions, and the
in-plane dimensions of the target were also
investigated. Results showed that the contact duration
INDIAN J. ENG. MATER. SCI., JUNE 2010
188
was higher for the plate with the smallest in-plane
dimensions, independently of the mass of the
impactor. Both contact force and the central
deflection increase by increasing the mass and
velocity of the impactor.
Freitas et al.22
have determined the mechanisms of
damage growth of the impacted composite laminates.
For this purpose, a series of impact tests were carried
out on carbon/epoxy composites. Two stacking
sequences and four different elastic behaviours with a
different number of interfaces were used. Results
showed that the delaminated area, due to impact
loading, depends on the number of interfaces between
plies. Two failure mechanisms were identified, which
are influenced by the stacking sequence and the
thickness of the plates. This study is focused on the
impact behaviour of glass/epoxy composite plates
with [0/±θ/90]S orientation, at equal energy (40 J),
equal velocity (2 m/s) and equal impactor mass (5
kg). Five different fiber directions and three different
plate thicknesses were selected to investigate the
stacking sequence and thickness effect. The
overlapped delamination area was also determined for
both parameters.
Manufacturing of the Composite Laminates
The unidirectional E-glass fabric having a weight
of 509 g/m2 was used as reinforcing material. CY225
epoxy and HY225 hardener were mixed at 90°C and
applied to the E-glass fabric. Then, it was cured under
250 kPa pressure for about 2 h at 120°C. The
composite plate consists of eight unidirectional layers.
At the end of the manufacturing process, the final
thickness of the composite plate with the
[0/30/60/90]S orientation was measured as about
2.9 mm. The mechanical properties of the composite
lamina used in this numerical study are given in
Table 123
.
Experimental Study The impact tests are performed at impact energy of
20 J to compare with the numerical results by using
Fractovis Plus impact test machine in the Composite
Research Laboratory of Dokuz Eylül University. The
Fractovis Plus impact test system is a test system
suitable for a wide variety of applications requiring
low to high impact energies under various
temperatures. The impactor with a hemispherical nose
of 12.7 mm in diameter is used. The testing machine
has a force transducer with capacity of 22.24 kN. The
total masses of the impactor are chosen as 5 kg and 10
kg (included impactor mass and crosshead mass). The
impact specimens with dimensions of 100 mm × 100
mm are clamped by using a pneumatic fixture.
Schematic diagram of the impact test machine is
shown in Fig. 1.
As the impactor dropped and approached the
composite specimen, its time trigger passed through a
time sensor right before contact-impact occurred. The
initial impact velocity was then calculated from the
distance between two edges on the time trigger and
the time interval they passes through the sensor. Once
impact began, the contact forces at many consecutive
instants were detected by the force transducer
attached to the impactor. The force history data was
recorded by data acquisition system (DAS). Data
points collected during a test are up to 16000 for each
channel. Sampling rate is 1 kHz-2 MHz for each
channel. Acceleration of the impactor is obtained by
dividing difference between impact force and total
weight of the impactor (gMtotal) to total weight of the
impactor deflection derives from a double integration
of acceleration of the impactor24
.
Table 1―The mechanical properties of the composite lamina23
Longitudinal modulus, E1 (GPa) 40.51
Transverse modulus, E2 (GPa) 13.96
In-plane shear modulus, G12 (GPa) 3.10
Poison’s ratio, ν12 0.22
Long. tensile strength, Xt (MPa) 783.30
Trans. tensile strength, Yt (MPa) 64.00
Long. comp. strength, Xc (MPa) 298.00
Trans. comp. strength, Yc (MPa) 124.00
Interlaminar shear strength, Si (MPa) 38.00
In-plane shear strength, S12 (MPa) 69.00
Fig. 1―Schematic illustration of the impact tester
KARAKUZU et al.: DAMAGE PREDICTION IN GLASS/EPOXY LAMINATES
189
Fig. 2―A basic model of impact problem
Numerical Study The impact problem can be described as shown in
Fig. 2. An impactor that has a mass of m and a radius
of r with the velocity of V drops on the center of the
composite plate that is fixed from four sides.
In a laminated composite plate; the impact damage
consists of delamination, matrix crack and fiber
fracture. Delamination is an inter-ply event and has a
direct relation with the differences in the ply
orientations. The delamination does not occur in any
inter-ply, which has the same fiber orientation angles.
Delamination area has a peanut-like shape, in which
the waist of the peanut is under the contact point. The
delamination occurs when the contact force reaches a
threshold value. This value could not be predefined
including all laminates or a specified orientation.
Based on the transversely impact condition, there
are only three stress components that can contribute to
the initial matrix cracking in the 90° layers. These are
the interlaminar shear stress σyz, in-plane tensile stress
σyy and out-of-plane normal stress σzz. However, the
out-of-plane normal stress is very small in comparison
with the other two stress components during the entire
impact event. Hence, the matrix cracking criterion can
be expressed as (σyy ≥ 0)25
,
2 2
21 failure
1 nofailure
n n
yy yz M
Mn n
Mt i
ee
eY S
σ σ ≥+ = <
… (1)
where the subscripts of x and y are the local
coordinates of the nth layer parallel and normal to the
fiber directions, respectively, and z is the out-of-plane
direction. n
tY and n
iS are in situ ply transverse tensile
strength and interlaminar transverse shear strength
within the nth ply of laminate under consideration,
respectively. n
yzσ and n
yyσ are the averaged
interlaminar transverse shear stress and the averaged
in-plane transverse tensile stress, respectively. For
determining any additional matrix failure, the
criterion must be applied to the other layers. If no
additional matrix cracking is found during impact,
then the impactor velocity associated with the first
matrix cracking is referred to as the impact velocity
threshold, which is the velocity required to just cause
the initial impact damage of the laminate.
Delamination can be initiated from the matrix
cracking in a ply within the laminate. Basically, there
are two types of critical cracks for initiating the
delamination resulting from impact as26
: one is the
shear crack and the other is the bending crack. For the
shear-crack, delamination can occur due to the
interlaminar longitudinal shear stress σxz in the layer
right below the interface and the interlaminar
transverse shear stress σyz in the layer right above the
interface. However, for the bending-crack,
delamination can occur due to the interlaminar
longitudinal shear stress σxz in the layer right below
the interface and the in-plane bending stress σyy in the
layer right below the interface. By taking into
consideration both failure mechanisms, delamination
occurs only when the following two sequential
conditions are met:
(i) One of the p1y groups intimately above or
below the concerned interface has failed due to
matrix cracking.
(ii) The combined stresses governing the
delamination growth mechanisms through the
thicknesses of the upper and lower ply groups of
the interface reaches a critical value.
Based on the above statement, impact-induced
delamination criterion for low velocity impact
proposed by Choi & Chang27
is
2 22 11
1 1
21 1
1 1
1 failure
1 nofailure
0
0
n nnyz yyxz
a n n n
i i
D
D
n nDt yy
n n
c yy
DS S Y
e
ee
Y Y if
Y Y if
σ σσ
σ
σ
++
+ +
+ +
+ +
+ +
≥
<=
= ≥ = <
… (2)
INDIAN J. ENG. MATER. SCI., JUNE 2010
190
where Da is a constant which has to be determined
from the experiments. The superscripts n and n+1
correspond to the upper and lower plies of the nth
interface, respectively. n
yzσ and 1n
yyσ+ are the
averaged interlaminar and in-plane transverse stresses
within the nth and ( 1)th
n + ply, respectively. 1n
xzσ
+ is
the averaged interlaminar longitudinal shear stress
within the ( 1)thn + ply.
In this study, 3DIMPACT code was used as a
solver. It is a ForTran-based transient dynamic finite
element analysis code, which can calculate the
stresses and contact forces according to composite
plates during impact event. It can also be used for
predicting the threshold of impact damage and
initiation of delaminations. An eight-point brick
element and the direct Gauss quadrature integration
scheme were used through the element thickness to
account for the change in material properties. The
Newmark scheme was adopted to perform time
integration from step-to-step. A contact law
incorporated with the Newton-Raphson method was
applied to calculate the contact force during impact.
The transient analysis has conducted in the laminates
based on three-dimensional linear elasticity. During
the analysis, the composite material is assumed as
homogeneous and orthotropic in each layer. The
3DIMPACT code allows the evaluation of
delamination areas by means of stress analysis and
above mentioned damage criteria.
The finite element mesh used for the calculation is
given in Fig. 3. The laminate is divided into N × M ×
Q elements as shown in this figure. The overall shape
is a square which has a dimension of 76.2 mm per
edge. A total of four elements are used through the
thickness of the laminate. Therefore, 576 (12×12×4)
elements were used in the numerical calculations for
generating the results.
Results and Discussion In this paper, the stacking sequence and thickness
effects on the impact behaviour of the glass/epoxy
composite plate was investigated numerically at equal
energy (40 J), equal velocity (2 m/s) and equal
impactor mass (5 kg). In this manner, five different
orientations as [0/±15/90]S, [0/±30/90]S, [0/±45/90]S,
[0/±60/90]S and [0/±75/90]S and three different plate
thicknesses as 2.9 mm, 5.8 mm and 8.7 mm were
chosen to investigate the stacking sequence and
thickness effects on glass/epoxy composite plates.
To verify the accuracy of the 3DIMPACT code,
numerical and experimental studies are carried out in
the glass/epoxy composite plate with the
[0/30/60/90]S orientation subjected to an impactor
with 20 J impact energy and 5 kg mass (2.828 m/s
velocity). The comparison between experimental and
numerical results in the same conditions is shown in
Fig. 4. It is seen from figure that the contact forces
obtained from both analyses are close to each other.
The delamination area for [0/30/60/90]S orientated
plates subjected to 20 J impact energy at 2 m/s
velocity (10 kg mass) is given in Fig. 5. To obtain the
overall delamination for numerical analysis, firstly,
six different delamination areas, which have been
located in different interfaces, are plotted. Afterwards,
all delaminations are added by overlapping them onto
each other. Thus, a good agreement is obtained
between the numerical analysis and the experimental
study.
After a good agreement between experimental and
numerical study is reached, the numerical analysis is
Fig. 3―Finite element model of the composite plate
5kg-20J
(2.828m/s)
0
2
4
6
8
10
0 1 2 3 4 5 6 7
Contact Time [ms]
Conta
ct
Forc
e [
kN
]
Experiment
3DIMPACT
Fig. 4―Contact force-contact time history for [0/30/60/90]S
laminate
KARAKUZU et al.: DAMAGE PREDICTION IN GLASS/EPOXY LAMINATES
191
carried out for various cases. The maximum contact
force and maximum deflection with different fiber
orientation at equal energy, equal mass and equal
velocity cases are given for 2.9 mm plate thickness in
Figs 6 and 7, respectively. It is seen from Fig. 6 that
the contact force increases by increasing ±θ for three
cases due to variation in greater orientation
differences. Thus, it can be said that the smaller ±θ
cause smaller contact forces. The maximum contact
force decreases with increasing impactor mass at
equal energy case, while it increases with increasing
impact energy at equal mass and also it increases by
increasing impactor mass at equal velocity for all
fiber orientations. The maximum contact force
decreases suddenly after a certain impactor mass.
After this level of impactor mass (5 kg), the
maximum contact force does not change with
impactor mass significantly. Thus, the lower impactor
mass with higher impact velocity has the greater
contact force (Fig. 6a).
However, this behaviour is different from the result
seen in Figs 6b and 6c. The lower impact energy with
lower impactor velocity cause the lower contact force
(Fig. 6b). In addition, the lower impactor mass with
lower impact energy has the smaller contact force
(Fig. 6c). It can be concluded from Fig. 6 that the
more effective impact parameter on the maximum
contact force is impact velocity.
The whole contrast in the behaviour of contact
force occurs for the maximum deflection; i.e., the
maximum deflection decreases by increasing ±θ at
equal energy, equal mass and equal velocity cases
(Fig. 7). The lower mass with higher velocity has the
higher deflection (Fig. 7a) while the lower energy
with lower velocity and the lower mass with lower
energy cause the lower deflection (Figs 7b and 7c).
To demonstrate the plate thickness effect on the
maximum contact force and the maximum deflection
are given in Figs 8 and 9 for the three cases explained
formerly. The thicker composite have more stiffness.
55 mm
55 mm
Interface 1 (0/30) Interface 2 (30/60) Interface 3 (60/90)
Interface 4 (90/60) Interface 5 (60/30) Interface 6 (30/0)
Overall-Numerical Overall-Experimental
Fig. 5―Delamination area for [0/30/60/90]S orientated plates subjected to 20 J impact energy at 2 m/s impact velocity
INDIAN J. ENG. MATER. SCI., JUNE 2010
192
0
2
4
6
8
10
12
14
5kg 10kg 15kg 20kg
Impactor mass [kg]
Max
imu
m c
on
tact
forc
e [
N] 1
15 30 45 60 75
0
2
4
6
8
10
12
14
10 J 20 J 30 J 40 J
Impact energy [J]
Max
imu
m c
on
tact
forc
e [
N] 1
15 30 45 60 75
0
2
4
6
8
10
12
14
5kg 10kg 15kg 20kg
Impactor mass [kg]
Max
imu
m c
on
tact
forc
e [
N] 1
15 30 45 60 75
Fig. 6―Maximum contact force variation with different
orientation for 2.9 mm thickness at (a) equal energy (40 J),
(b) equal mass (5 kg) and (c) equal velocity (2 m/s)
0
2
4
6
8
10
5kg 10kg 15kg 20kg
Impactor mass [kg]
Max
imu
m d
efl
ecti
on
[m
m]
12
15 30 45 60 75
0
2
4
6
8
10
10 J 20 J 30 J 40 J
Impact energy [J]
Max
imu
m d
efl
ecti
on
[m
m]
12
15 30 45 60 75
0
2
4
6
8
10
5kg 10kg 15kg 20kg
Impactor mass [kg]
Max
imu
m d
efl
ecti
on
[m
m]
12
15 30 45 60 75
Fig. 7―Maximum deflection variations with different orientation
for 2.9 mm thickness at (a) equal energy (40 J), (b) equal mass
(5 kg) and (c) equal velocity (2 m/s)
(a)
(b)
(c)
(a)
(b)
(c)
KARAKUZU et al.: DAMAGE PREDICTION IN GLASS/EPOXY LAMINATES
193
Hence, higher plate thicknesses cause higher contact
force and lower deflection. But contact force and
deflection rates decrease by the increasing plate
thickness.
The overall delamination areas in [0/±45/90]S
stacking sequences are given in Fig. 10 for equal
energy, equal mass and equal velocity to investigate
the delamination behaviour of the glass/epoxy
laminated composite plates. It can be seen from the
figure that the overall delamination does not change
significantly by the impactor mass at the equal energy
case. However, for the case of 40 J - 5 kg, it is larger
then the other impactor masses due to the highest
impactor velocity (Fig. 10a). In the case of equal
impactor mass, the overall delamination area
increases by the increasing impact energy (Fig. 10b).
It also increases by increasing the impactor mass
while the impactor velocity is kept constant
(Fig. 10c). Figure 11 demonstrates the overall
delaminations for different orientations at 20 J impact
energy and 5 kg impactor mass. From the figure, it
can be concluded that the overall delamination does
0
5
10
15
20
25
30
35
40
2,9 mm 5,8 mm 8,7 mm
Plate thickness [mm]
Max
imu
m c
on
tact
forc
e [
N] 1
15 30 45 60 75
0
5
10
15
20
25
30
35
40
2,9 mm 5,8 mm 8,7 mm
Plate thickness [mm]
Max
imu
m c
on
tact
forc
e [N
] 1
15 30 45 60 75
Fig. 8―Variation of maximum contact force by plate thickness at (a) 5 kg mass and 4 m/s velocity (40 J) and (b) 5 kg mass and
2 m/s velocity (10 J)
0
2
4
6
8
10
2,9 mm 5,8 mm 8,7 mm
Plate thickness [mm]
Max
imu
m d
efl
ecti
on
[m
m] 1 15 30 45 60 75
0
2
4
6
8
10
2,9 mm 5,8 mm 8,7 mm
Plate thickness [mm]
Max
imu
m d
efl
ecti
on
[m
m] 1
15 30 45 60 75
Fig. 9―Variation of maximum deflection by plate thickness at (a) 5 kg mass and 4 m/s velocity (40 J) and (b) 5 kg mass and
2 m/s velocity (10 J)
(a)
(a)
(b)
(b)
INDIAN J. ENG. MATER. SCI., JUNE 2010
194
40J-5kg
5kg-10J
2 m/s-5kg
40J-10kg
5kg-20J
2 m/s-10kg
40J-15kg
5kg- 30J
2m/s-15kg
40J-20kg
5kg- 40J
2 m/s-20kg
(a) (b) (c)
Fig. 10―Delamination areas for [0/±45/90]S at (a) equal energy (40 J), (b) equal mass (5 kg) and (c) equal velocity (2 m/s)
KARAKUZU et al.: DAMAGE PREDICTION IN GLASS/EPOXY LAMINATES
195
[0/±15/90]S
[0/±30/90]S
[0/±45/90]S
[0/±60/90]S
[0/±75/90]S
Fig. 11―Overall delaminations for different orientation at 20 J impact energy and 5 kg impactor mass
0
500
1000
1500
2000
15 30 45 60 75
Fiber orientation ( o
)
Del
amin
atio
n a
rea
(mm
2
) t=2.9 mm
t=5.8 mm
t=8.7 mm
0
500
1000
1500
2000
15 30 45 60 75
Fiber orientation ( o
)
Del
amin
atio
n a
rea
(mm
2
) t=2.9 mm
t=5.8 mm
t=8.7 mm
(a) (b)
0
500
1000
1500
2000
15 30 45 60 75
Fiber orientation ( o
)
Del
amin
atio
n a
rea
(mm
2
) t=2.9 mm
t=5.8 mm
t=8.7 mm
0
500
1000
1500
2000
15 30 45 60 75
Fiber orientation ( o
)
Del
amin
atio
n a
rea
(mm
2
) t=2.9 mm
t=5.8 mm
t=8.7 mm
(c) (d)
Fig. 12―Overall delamination areas at equal energy (40 J) for (a) 5 kg, (b) 10 kg, (c) 15 kg and (d) 20 kg
INDIAN J. ENG. MATER. SCI., JUNE 2010
196
not significantly change with fiber orientation.
However, the delamination zone orientates along the
±θ fiber orientation.
To compare the thickness effect on the
delamination area for different orientation angles, the
overlapped delamination areas are given in Figs 12-14
for equal energy, equal mass and equal velocity,
respectively. It is seen from Fig. 12 that the overall
delamination area decreases by increasing the plate
thickness because of the stiffer laminate. In addition,
the bending cracks cannot occur. The inertia effect
also causes an increase in the maximum contact force.
Hence, the delamination decreases by increasing of
the inertia. In addition, the delamination decreases by
the increase in ±θ. But it does not change significantly
by increasing the impactor mass at the same impact
energy (Fig. 12). For the equal mass case, the overall
delamination increases by increasing the impact
energy and decreases by increase in ±θ (Fig. 13).
Delamination areas for equal impact velocity in Fig.
14 show similar behaviour to ones in Fig. 13.
Conclusions In this study, the impact behaviour of the
glass/epoxy composite plate with [0/±15/90]S,
[0/±30/90]S, [0/±45/90]S, [0/±60/90]S and [0/±75/90]S
fiber orientations were investigated numerically at
equal energy (40 J), equal velocity (2 m/s) and equal
impactor mass (5 kg). The following conclusions can
be drawn from the results obtained:
0
500
1000
1500
2000
15 30 45 60 75
Fiber orientation ( o
)
Del
amin
atio
n a
rea
(mm
2
) t=2.9 mm
t=5.8 mm
t=8.7 mm
0
500
1000
1500
2000
15 30 45 60 75
Fiber orientation ( o
)
Del
amin
atio
n a
rea
(mm
2
) t=2.9 mm
t=5.8 mm
t=8.7 mm
(a) (b)
0
500
1000
1500
2000
15 30 45 60 75
Fiber orientation ( o
)
Del
amin
atio
n a
rea
(mm
2
) t=2.9 mm
t=5.8 mm
t=8.7 mm
0
500
1000
1500
2000
15 30 45 60 75
Fiber orientation ( o
)
Del
amin
atio
n a
rea
(mm
2) t=2.9 mm
t=5.8 mm
t=8.7 mm
(c) (d)
Fig. 13―Overall delamination areas at equal impactor mass (5 kg) for (a) 10 J, (b) 20 J, (c) 30J and (d) 40 J
KARAKUZU et al.: DAMAGE PREDICTION IN GLASS/EPOXY LAMINATES
197
(i) The results obtained from numerical analysis are
close to the results obtained from the
experimental study.
(ii) The contact force increases by increasing ±θ at
equal energy, equal mass and equal velocity.
(iii) The lower impactor mass with higher impact
velocity causes greater contact forces. However,
the lower impact energy with lower impact
velocity and lower impact energy with lower
impactor mass cause lower contact forces.
(iv) The lower mass with higher velocity causes
higher deflection while the lower energy with
lower velocity and lower mass with lower
energy cause lower deflection.
(v) Higher plate thicknesses cause higher contact
force and lower deflection. But contact force and
deflection rates decrease by increasing the plate
thickness.
(vi) The overall delamination area increases by the
increasing impact energy. However, it does not
significantly change by increasing the fiber
orientation.
(vii) The overall delamination area decreases by
increasing the plate thickness.
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0
500
1000
1500
2000
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)
Del
amin
atio
n a
rea
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2
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t=5.8 mm
t=8.7 mm
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Fiber orientation ( o
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atio
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