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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


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)


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


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


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)


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


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


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


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)


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)


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


)

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


)

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


)

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


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


)

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


)

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


)

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


)

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


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

15 30 45 60 75


)

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


)

Del

amin

atio

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rea

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t=5.8 mm

t=8.7 mm

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0

500

1000

1500

2000

15 30 45 60 75


)

Del

amin

atio

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rea

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t=5.8 mm

t=8.7 mm

0

500

1000

1500

2000

15 30 45 60 75


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Del

amin

atio

n a

rea

(mm

2) t=2.9 mm

t=5.8 mm

t=8.7 mm

(c) (d)

Fig. 14―Overall delamination areas at equal impact velocity (2 m/s) for (a) 10 J, (b) 20 J, (c) 30 J and (d) 40 J


198

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