international journal impact
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www.elsevier.com/locate/ijimpeng
Author’s Accepted Manuscript
Repeated impact response of hand lay-up and vacuum
infusion thick glass reinforced laminates
Giovanni Belingardi, Maria Pia Cavatorta, Davide
Salvatore Paolino
PII: S0734-743X(07)00029-2
DOI: doi:10.1016/j.ijimpeng.2007.02.005
Reference: IE 1467
To appear in: International Journal of Impact
Received date: 29 September 2006
Revised date: 7 December 2006
Accepted date: 28 February 2007
Cite this article as: Giovanni Belingardi, Maria Pia Cavatorta and Davide Salvatore Paolino,
Repeated impact response of hand lay-up and vacuum infusion thick glass reinforced
laminates, International Journal of Impact (2007), doi:10.1016/j.ijimpeng.2007.02.005
This is a PDF file of an unedited manuscript that has been accepted for publication. As
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REPEATED IMPACT RESPONSE
OF HAND LAY-UP AND VACUUM INFUSION
THICK GLASS REINFORCED LAMINATES
Giovanni Belingardi, Maria Pia Cavatorta∗, Davide Salvatore Paolino
Mechanical Engineering Department – Politecnico di Torino
Corso Duca degli Abruzzi, 24 – 10129 Torino (Italy)
ABSTRACT
Vacuum infusion (VI) is being considered as a viable alternative to more traditional
hand lay-up (HL). Main reason in favor of the more costly technique is the cleaner and
friendlier work environment. Moreover, VI potentially offers another important benefit
over HL in that prepreg levels of resin may be achieved, resulting in stronger and lighter
laminates. The present paper compares the two manufacturing techniques on the basis
of the response to repeated impact loading. The laminate is a thick non-symmetric glass
fiber reinforced plastics intended for nautical application. Four impact velocities (1.5m/s, 2.2 m/s, 3.1 m/s and 3.8 m/s) were considered, and a minimum of four specimens
for any given velocity were subjected to forty repeated impacts or up to perforation. The
impact response was evaluated in terms of damage progression by visual observation of
the impacted specimens, evolution of the peak force and of the bending stiffness with
the number of impacts and by calculating the Damage Index (DI), a damage variable
recently proposed by the authors to monitor the penetration process in thick laminates.
Results point out that for impact velocities for which no perforation occurs within test
duration, the experimental data essentially overlap. On the contrary, for perforation
tests, HL specimens survived more impacts before perforating absorbing more total
energy than VI specimens. Plots of the DI variable against the number of impacts were
observed to exhibit an initial linear portion, owing to a stable process of damage
accumulation within the laminate, and to undergo an unstable growth a few impacts
before perforation. When comparing the VI and HL specimens it was observed that,
given an impact energy, the level of damage at first impact as well as the rate of stable
damage accumulation is alike for the two sets of specimens. On the contrary, it is the
number of impacts of the stable damage accumulation region which is lower for VI
specimens.
KEYWORDS: low velocity impact, damage accumulation, glass fiber reinforced
composite, nautical applications.
INTRODUCTION
In the nautical sector, where large products are involved and where there is no massive
series production, Hand Lay-up (HL) is by large the most widely used manufacturing
∗ Corresponding author. Fax number: +39.011.5646999. e-mail address: [email protected]
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technique. However, in the HL technique, a mayor health risk for operators comes from
the resin fumes that exhale from the open mould on which the operator needs to hover
over while saturating the laminate by hand. Producers of fiber reinforced composite
laminates are thus looking for cleaner and friendlier work environment manufacturing
processes. In this respect, Vacuum Infusion (VI) is now being considered as a viable
alternative to the more traditional HL technique.
While in a typical HL, reinforcements are laid into the mould and manually wet out
using brushes or rollers and then vacuum is used to remove the resin in excess, VI takes
a different approach, in that a vacuum is drawn while the materials are still dry. Once
vacuum is achieved, resin is literally sucked into the laminate via carefully placed
tubing. Ideally, any excess resin that is introduced will eventually be sucked out into the
vacuum line.
Therefore, besides environmental issues, VI potentially offers another important benefit
over HL, in that it should allow for a very predictable resin usage approaching prepreg
levels of resin content. Because of the improvement in the fiber-to-resin ratio, laminates
manufactured by VI should be stronger and lighter as compared to laminates
manufactured by HL.
The paper presents data of a comparative experimental study conducted on a glass fiber
reinforced plastics for nautical application, manufactured by HL and VI. Comparison is
made on the laminate response to low velocity repeated impacts, a loading condition of
particular relevance for naval or nautical applications. Recently few papers have
addressed the problem of repeated impact loading [1-5.]
Low velocity impacts on laminates are known to significantly reduce the laminate
strength and stiffness, mainly as a consequence of multiple stacked delaminations that
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are produced by the impact loading at a number of interfaces through the thickness of
the composite laminate. In [1,2], typical SN curves for glass fiber reinforced plastics are
drawn. In [1], an endurance limit above 104
impacts was observed. As it is the case for
classical fatigue, the endurance limit corresponds to a condition of no degradation of the
composite mechanical characteristics. In [2], SEM observations revealed that, even
without any visible external damage, microcracks in the resin can produce internal
delaminations, thus reducing the composite laminate strength. For composites, the
damage induced by impact loading is more subtle than in metals, as it is often not
detectable, beginning on the non-impacted surface or in the form of internal
delamination. Matrix cracking is chronologically the first damaging mode.
Delaminations are initiated by critical matrix cracks that are in fact emerging at the
interfaces between layers. Curves of damage evolution against the number of impacts
reveal three regions at various energy levels. At low impact energies, the damage
process is governed by initiation and multiplication of delamination, with the non-
impacted face of the specimen being the first region to delaminate. As the impact
energy increases, saturation of delamination is achieved. Beyond saturation of the
delamination process, there is an acceleration in the damage accumulation until final
failure, which occurs by ply cracking with fiber breakage. In [3], tensile and
compressive static tests were performed on beam-like specimens cut from carbon/epoxy
plates subjected to repeated impact tests. Test results point out that, for tests in which
perforation is not achieved within test duration, a degradation in residual properties
occurs only for specimens that include the damage impact zone. Impacts at higher
energy levels induce more damage (i.e. lower residual compressive and tensile strength)
than a number of lighter impacts. On the contrary, for tests in which perforation is
achieved, the degradation of residual properties is the same regardless of the energy per
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single impact. Moreover, the damage sustained in the area surrounding the contact point
is higher for a laminate perforated in several impacts with less energy per impact.
To the authors’ best knowledge all available experimental papers on repeated impact
tests are for thin laminates which rarely find application in the marine industry. As
correctly pointed out in [6], the advancement from thin to thick laminates is not trivial,
as thick composite laminates behave quite differently from their thin counterparts. One
important aspect differentiating thick from thin laminates is the extent of the penetration
process which obliges to a clear distinction between laminate penetration and
perforation. For thick laminates, the energy required by the dart to go through the
laminate can not be neglected.
In the present paper a thick glass reinforced plastics laminate is tested to repeated
impacts, with an emphasis on the comparison between the more traditional HL
manufacturing technique and a viable alternative such as VI. Four falling heights are
considered, corresponding to conditions of no perforation within test duration and to
conditions of laminate perforation. The impact response is evaluated in terms of damage
progression by visual observation of the impacted specimens, evolution of the peak
force (maximum of the load-displacement curve) and of stiffness loss as a function of
impact number, and by calculating the Damage Index (DI), a damage variable recently
proposed by the authors [7] to monitor the penetration process in thick laminates.
THE DAMAGE INDEX
In [8,9] Belingardi & Vadori introduced the Damage Degree (DD) to account for
damage accumulation in thin laminates.
Defined as the ratio between the absorbed energy Ea and the impact energy Ei (Figure
1):
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≡=
n penetratioafter E
E
n penetratiotoup E
E
DD
sa
a
i
a
,1
,
(1)
the DD was shown to increase rather linearly with the impact energy to reach the value
of one at laminate penetration. A saturation energy level Esa was defined as the impact
energy at which the DD regression curve reaches the value of one. This energy
threshold is of practical and theoretical interest since it defines the maximum energy
level the laminate can dissipate with no penetration and by means of internal damage
mechanisms only [9].
As pointed out in [6], while a single energy threshold is generally sufficient to define
the impact characteristics of thin laminates, for thick laminates two different threshold
values are to be defined: the penetration threshold Pn and the perforation threshold Pr .
The penetration threshold is identified at the first time the absorbed energy reaches the
level of impact energy and it is therefore conceptually the same as the saturation energy.
For impact energies above the penetration threshold, the impactor moves deeper into the
laminate. Once the impact energy is high enough, perforation eventually takes place.
The impact energy is higher than the absorbed energy and the energy in excess is
retained in the impactor for post-perforation motions. Between the penetration and
perforation thresholds, there exists a range, named by Liu “the range of the penetration
process”, in which the impact energy and the absorbed energy are equal to each other
but which represent different stages of the penetration process with the impactor
moving deeper and deeper into the specimen as the impact energy increases.
However, by its definition the DD is unable to differentiate between conditions of
laminate penetration and perforation since over the entire penetration process the
absorbed energy would be equal to the impact energy and the DD equal to one. To
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account for how far the impactor has moved down into the laminate, the maximum
deflection sMAX registered during the impact test was introduced into definition of the
DI:
QS
MAX
s s DD DI = (2)
Normalization by the maximum stroke of the quasi-static penetration test sQS allows for
a non-dimensional quantity. The sQS value as well as the sMAX value in perforation tests
correspond to the stroke value at which the force becomes constant (Figure 2) and equal
to the friction force of the dart sliding through the penetrated specimen. For all the
laminates tested in [7] the sMAX measured in perforation tests was practically equal to
sQS; so that the DI reaches the value of one at laminate perforation.
EXPERIMENTAL
The laminate under study is a glass fiber reinforced plastics. The stacking sequence is
reported in Table 1. Two resin systems are used: vinylester and polyester. Main reason
for the asymmetry of the laminate, in both the resin system and the stacking sequence, is
cost reduction. The interface between the two resins is located at about two thirds of the
laminate thickness from the vinylester face. The fiber reinforcement varies from 35% in
weight in the mat, 45% in the bidirectional lamina and 50% in the unidirectional lamina.
With the purpose of investigating the possibility of replacing the current manufacturing
technique, i.e. HL, with a more environment friendly technique such as VI, static
indentation and impact tests were performed on two sets of specimens manufactured by
the two technologies. Specimens were cut from plates. The thickness of the plates was
measured at several locations showing reduced scatter. However, for both technologies
the achieved thickness values were quite distant from the design protocol. The average
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thickness for HL was 10.13 mm (standard deviation: 0.25 mm), while for VI was 9.72
mm (standard deviation 0.05 mm).
For both sets, specimens were impacted on the polyester face of the laminate. Previous
tests performed on HL specimens [10] have indeed shown that, while higher peak forces
are achieved by impacting the vinylester face, the number of impacts to perforation is
greater when impacting the polyester face. The experimental result was explained by
considering that, in the case it is the polyester face to be impacted, the more ductile
resin –vinylester– is on the rear target face where delamination is known to initiate [2],
thus limiting the progression of damage. Moreover, it was observed that the condition
of laminate perforation is associated with the splitting of laminae, which always occurs
at the interface between the two resin systems (Figure 3). Therefore perforation is
delayed in the case the impacted face is polyester, being the interface between the two
resin systems more distant from the rear target face. For repeated impact tests, four
falling heights were considered (125 mm, 250 mm, 500 mm, 750 mm) corresponding to
four impact velocities (1.5 m/s, 2.2 m/s, 3.1 m/s and 3.8 m/s) and a minimum of four
specimens for any given velocity were subject to forty repeated impacts or up to
perforation. Single impact tests were also performed at 4.4 and 6.3 m/s (corresponding
to the maximum height of the drop-dart apparatus) to investigate possible strain-rate
effects.
Impact tests were performed according to ASTM 3029 standard [11] using an
instrumented free-fall drop dart testing machine. The impactor has a total mass of 20 kg,
its head is hemispherical with a radius of 10 mm. Stainless steel was chosen for its high
hardness and resistance to corrosion. The maximum falling height of the testing
machine is 2 m, which corresponds to a maximum impact energy of 392 J. The drop-
weight apparatus was equipped with a motorized lifting track. The collected data were
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stored after each impact and the impactor was returned to its original starting height.
Using this technique, the chosen impact velocity was consistently obtained in
successive impacts. Because, the target holder was rigidly attached to the frame of the
testing device, the tup struck the specimen each time at the same location. By means of
a piezoelectric load cell, force-time curves were acquired. The acceleration history was
calculated dividing the force term by the impactor mass. The displacement was obtained
by double integration of the acceleration and thus force-displacement curves were
plotted. By integration of the force-displacement curves, deformation energy-
displacement curves were then obtained. Initial conditions were given with the time axis
having its origin at the time of impact. At time t=0, the dart coordinate is zero and its
initial velocity can be obtained by the well known relationship:
h g v ∆= 20 (3)
where ∆h is defined as the height loss of the center of mass of the dart with respect to
the reference surface [8]. The drop dart machine used in the study is equipped with an
optoelectronic device for measurement of the impact velocity. Agreement between
measured and theoretical values was very good.
Square specimen panels, with 100 mm edge, were clamped with a 76.2 mm inner
diameter, and fixed to a rigid base to prevent slippage of the specimen (Figure 4). The
clamping system makes use of pre-loaded springs to provide an adequate and repeatable
uniform pressure all over the clamping area.
Prior to impact tests, a series of static indentation tests were performed to get
information on the material stiffness and strength characteristics, which serves as a
starting point to decide on the falling heights of impact tests. For quasi-static
penetration tests, specimens were tested using a servo-hydraulic machine with
maximum loading capacity of 100 kN. The hydraulic actuator was electronically
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controlled in order to perform constant velocity tests. Signals of the force applied by the
actuator and of the actuator stroke were acquired in time with an appropriate sampling
rate.
To assess the progression of damage, for one series of repeated impact tests at any given
impact velocity, pictures of the specimen impacted and rear face as well as of the
specimen thickness were taken after each impact. A software program was used to
determine the area of delamination, readily seen under light due to a change in the
opacity of the resin. In addition to visual observation of impacted specimens, for all
tests variation of bending stiffness against impact number was evaluated from the force-
displacement curves.
TESTS AND RESULTS
Figure 5a depicts representative force-displacement curves obtained in quasi-static
perforation tests for the two sets of laminates. As it can be observed from the graph, the
first damage force for VI specimens is around 8 kN. The small plateau in the force
associated to the first damage was consistently observed in all tests performed on VI
specimens. For HL, the first damage force is around 10kN and leads to a variation in the
curve slope rather than to a plateau. While the first damage force appears to be slightly
higher for HL than for VI specimens, values for the maximum force are alike. As for the
laminate stiffness, values were calculated from the initial portion of the force-
displacement curves before the first damage takes place and are reported in Figure 5b.
The laminate stiffness was consistently higher for HL specimens. Over a minimum of
three tests for each manufacturing technology, calculated average values for the
laminate stiffness were 6493 N/mm for HL and 5736 N/mm for VI specimens.
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In analyzing the static test results, it should be reminded that the VI specimens are about
four tenths of a millimeter thinner than HL specimens. In this respect, it is interesting to
note that the different values obtained for the laminate stiffness are well explained by
the difference in thickness. Indeed, the ratio between the average stiffness values
(6493/5736= 1.13) is equal to the ratio of the thickness elevated to the third power
(10.13/9.72=1.04; (1.04)3= 1.13). Due to the thickness of the laminate under study, it is
the flexural behavior that affects the first damage event rather than the membrane
behavior as observed in very thin laminates [9]. Looking for a relationship between the
first damage force and the thickness is more troublesome. Indeed, for the case under
study, the difference in thickness is so limited that it would be speculative to comment
on a relationship between load and thickness elevated to the power of 3/2, as observed
by many authors with reference to the critical load in impact tests [12]. What can be
said is that for the first damage force (10/8=1.25) the manufacturing technology seems
to play a role in determining the value of the force and the type of damage involved.
Figures 6 show the force-displacement curves for two series of ten repeated impact tests
at 3.1 m/s run on HL and VI specimens. From the two graphs, the Damage Threshold
Load (DTL) [12] can be evaluated and compared for the two sets of specimens. A
correlation between the DTL and the laminate thickness is noticeable (12.8/11.6= 1.10).
The two series of curves depicted in Figure 6 clearly show that the laminate stiffness
diminishes impact after impact and that the highest reduction is achieved in the first few
impacts. The statement is supported by data shown in Figure 7, where the stiffness is
plotted against the impact number for the considered four impact velocities and for HL
and VI specimens. Reasonably, the total loss in laminate stiffness is greater for higher
impact velocities. Data observation points out that for VI specimens (solid symbols) the
stiffness loss is in essence concentrated in the first impact while the loss appears less
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abrupt for HL specimens (hollow symbols). However, when looking at the value of the
stiffness at perforation or the asymptotic value of no perforation tests, they are alike for
VI and HL specimens meaning that the total stiffness loss is greater for HL specimens.
Data points on the vertical axis show that the laminate stiffness is not constant at
different impact velocities. To investigate the laminate strain-rate sensitivity, additional
single impact tests were performed at 4.4 and 6.3 m/s (corresponding to the maximum
height of the drop-dart apparatus). Figures 8a and 8b plot average values for the
laminate stiffness and the initial damage load in the form of DTL, respectively, against
the impact velocity. Data obtained in quasi-static penetration tests are reported on the
vertical axis. The maximum available impact energy was not sufficient to perforate the
laminate in single impact tests, therefore no comparison could be carried out in terms of
maximum load. Although the investigated range of impact velocities is rather limited,
data plotted in Figure 8 demonstrate a strain-rate dependency of the stiffness and initial
damage force. Values of R 2 show a good fitness of a linear function on the experimental
data. The observed strain-rate dependency is likely to be owed to the large amount of
resin in the laminate.
In Figures 9 and 10 plots of the peak force and of the damage variable DI against the
impact number are presented. To avoid confusion among the experimental data,
separate graphs have been edited grouping data obtained in no perforation tests (Figures
9) and in perforation tests (Figures 10). For Figure 10 reported data are not average
values but refer to one series of repeated impact tests. Indeed, in perforation tests, the
number of impacts to perforation among different specimens may change by one or two
unities, thus causing scatter in the peak force and DI values in the one-two impacts
before perforation. On the contrary, for no perforation tests single test and average
values agree very well.
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Experimental results presented in Figures 9 show that, for no perforation tests, data for
the two sets of specimens are rather similar. A distinction should be made between the
two impact velocities. At the lowest velocity, for which both the peak force and the DI
reach a stable value indicating the reaching of a steady-state condition, data for the HL
and VI basically overlap. For the graphs at 2.2 m/s, in which the peak force continues to
slowly decrease impact after impact while the DI increases, suggesting a slow but
steady accumulation of damage, VI specimens appear to damage more than HL.
The result is confirmed in perforation tests (Figures 10). Main result that can be evinced
from Figure 10b is that the number of impacts to perforation is smaller for VI
specimens. The depicted difference of four impacts between VI and HL is the average
observed gap. When looking at the result, it is important to remind that the VI
specimens are about 4 tenths of a millimeter thinner than the HL (9.72 mm against
10.13 mm). However, taking into account the difference in thickness to give reason for
a difference in the number of impacts to failure is not straightforward, as it was for the
value of the stiffness. A possibility could be comparing the two sets of specimens on the
basis of the total impact energy density at perforation, that is on the basis of the product
between the number of impacts to perforation and the impact energy density per single
event (impact energy divided by the specimen volume). As it can be observed from
Figure 11, the total impact energy at perforation is higher for HL specimens, which
seem to exhibit a greater damage tolerance than VI. From Figure 11 it can also be noted
that, regardless of manufacturing technology, perforation is not associated with a
specific level of total energy. Higher impact energies per single event are more
damaging than a number of lighter impacts.
As for the graphs of the peak force (Figures 9a and 10a), it is interesting to note that for
all considered impact velocities, the maximum of the peak force is not reached in the
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first impact. The literature suggests two explanations for the phenomenon. In a series of
repeated impact tests run on carbon/epoxy composite laminate, Wyrick and Adams [3]
commented the initial increase in the peak force as the result of the compaction process
of the thin layer of unreinforced resin at the impacted surface. At low impact energy
levels, damage to the fibers near the surface is minimal and the compaction process
provides a harder surface for the next impact. In a series of single impact tests run at
different impact velocities on glass/epoxy laminates, Liu [6] observed that, even if
delamination takes place very early in the impact event, indentation and local matrix
cracking are the dominant damage modes up to the maximum peak force. The
maximum in the peak force therefore signals a turning point in the dominant damage
mode. In particular, up to maximum peak force the dominant damage modes are
indentation and local matrix cracking around the impacted region; while after being
loaded by the maximum peak force, the damage accumulation process is dominated by
delamination. For impact velocities above the maximum peak force, delamination
becomes the dominant damage mode while matrix cracking, and hence lamina splitting,
continue to grow.
When considering the curves depicted in Figures 9a and 10a, the compaction process
suggested in [3] may well explain the reaching of a steady-state value in the peak force
for the 1.5 m/s tests as well as the initial rise in the peak force observed for the 2.2 m/s
tests (Figure 9a). On the contrary, both visual observation of the impacted specimens
and the laminate stiffness loss (Figure 7) do not seem to support Liu’s findings. Indeed,
for the 3.1 and 3.8 m/s impact velocities, the stiffness loss is concentrated in the first
few impacts; moreover, visual observation of the impacted specimens did not reveal
anything noteworthy in the impacts before and after the maximum of the peak force. In
the authors’ opinion, it should be considered that the value of the peak force is affected
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by two opposite mechanisms: on one side the compaction of the resin which provides a
harder surface for the next impact event and on the other side the progression of damage
which reduces the laminate mechanical properties. Initially, the beneficial local surface
effect of resin compaction can determine an increase in the peak force, even if the
global bending stiffness of the laminate has decreased.
Figures 9a and 10a show that at increasing impact velocities, the impact number at
which the maximum of the peak force occurs generally decreases. Also, the decrease in
the peak force after the maximum value becomes sharper. Interestingly, while for
impact velocities causing no perforation the maximum value of the peak force increases
for increasing impact velocity (Figure 9a), in tests where perforation takes place the
maximum in the peak force is alike regardless of the test impact velocity (Figure 10a).
A few comments are worthwhile making on the DI plots of Figures 9b and 10b. In no
perforation tests (Figure 9b), the DI value is always very small. While for the 1.5 m/s
tests it remains constant at the value of 0.1 impact after impact, for the 2.2 m/s tests the
DI increases quite linearly with the impact number, owing to a slow but steady
accumulation of damage. DI values for VI specimens are slightly above those for HL
specimens. In perforation tests (Figure 10b), the DI shows an initial linear region
followed by an unstable growth a few impacts before perforation. During the
penetration process the DI growth is highly non linear. A value of one is reached at
complete laminate perforation. By looking at the DI values plotted in Figure 10b, it is
worth commenting that, given the impact velocity, the initial linear region of the curves
is alike for HL and VI specimens, both in terms of point values and of curve’s slope. In
other words, both the level of damage at first impact and the rate of stable damage
accumulation is the same regardless of the manufacturing technology. On the contrary,
it is the number of impacts sustained by the specimen before the onset of unstable
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damage growth which is smaller for VI specimens. As it can be evinced from Figure
10b, the number of impacts in the unstable region does not change between VI and HL
specimens. In [13], it was shown that a strong interfacial bond between the fiber and
resin matrix delays the initiation of severe damage modes, thus improving the overall
resistance to repeated impacts. Similarly, the reduced damage tolerance of VI specimens
could be ascribed to the larger amount of resin of HL specimens and to the fact that
manual wetting out of the prepregs assures good quality and homogeneous fiber
impregnation. One last comment is to be made regarding the slope of the DI vs. impact
number curves at different impact velocities. By looking at Figures 9b-10b, it is
apparent that the slope increases for increasing impact velocity, i.e. increasing impact
energy. Figure 12 plots values of the slope of the DI vs. impact number curves as a
function of the impact energy. A quadratic regression curve well fits the experimental
data. As already pointed out, no significant difference is found between HL and VI
specimens.
Figures 13 and 14 reproduce photographs of the laminate impacted and rear faces after
the first impact as well as after the impact prior to perforation, for HL and VI specimens
respectively. The impact velocity is 3.1 m/s. As it is noticeable from the pictures, the
damage appearance for the two laminates is rather different, in particular for what
concerns the extent of damage visible from the rear face after the first impact.
The extent of the delamination area in the laminate was assessed by the change in the
opacity of the resin. Clearly such evaluation was limited to the laminate faces and
neglects internal delaminations. Observation of the specimen thickness after each
impact event allowed to observe that prior to perforation no significant laminae splitting
occurs, while at perforation splitting of laminae takes place at the interface between the
two resin systems (Figure 3).
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By observing the two laminate faces, it can be noticed that in the case of VI specimens
(Figure 13), delamination extends to the whole unsupported region of the specimen
from the first impact. On the contrary, for HL specimens (Figure 14) the amount of
delamination in the rear face after the first impact is rather limited. Delamination
continues to grow as the number of impacts increases but does not appear to saturate the
whole specimen unsupported region even at perforation. The dissimilarity in the extent
of the delamination area visible from the rear face may explain the previously observed
difference in stiffness loss at first impact, i.e. the fact that for VI specimens the stiffness
essentially drops off in the first impact while it decreases less abruptly in HL specimens
(Figure 7). If the damage associated with delamination appears earlier in time for VI
specimens, the energy dissipation mechanism of fiber pull-out is on the contrary more
readily seen in HL specimens. By looking at the rear face of the two specimens it can be
noted that, while in VI specimens fiber damage is first limited to the contact area with
the impactor tup, for HL specimens fiber pull-out extends along the fiber directions
from the first impact.
CONCLUSIONS
The impact characteristics of a non-symmetric thick glass fiber reinforced plastics
laminate were investigated. To look into the possibility of manufacturing the laminate
through VI, two series of repeated impact tests up to forty impacts or to perforation
were run at different impact velocities, comparing VI with the more traditional HL
technology. The impact response was evaluated in terms of damage progression by
visual observation of the impacted specimens, evolution of the peak force and of the DI
with the number of impacts as well as through assessment of laminate stiffness
reduction.
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The following statements summarize the experimental results:
• no significant differences exist in the force, energy curves and damage
parameter for tests in which no perforation occurs;
• when the impact energy is such to determine laminate perforation, HL
specimens survived more impacts before perforating absorbing more total
energy;
• visual observation of the impacted specimens shows that in both series of tests
and for all the considered impact velocities, the growth of the delamination area
is concentrated in the first impacts. In particular, for VI specimens, the
delamination area saturates from the first impact, following closely the behavior
of the bending stiffness against the number of impacts;
• for both sets of specimens, perforation of the laminate is associated to laminae
splitting at the interface between the two resin systems;
• the damage variable DI is observed to initially grow linearly impact after impact
owing to a stable accumulation of damage to then undergo an abrupt growth a
few impacts before perforation. During the penetration process the DI growth is
highly non linear. A value of one is reached at complete laminate perforation;
• when calculating the rate of stable damage accumulation (slope of the DI vs.
impact number curves) in the initial linear portion, a quadratic relationship is
found between the rate of damage accumulation and the impact energy;
• given the impact energy, no significant difference in the rate of stable damage
accumulation is observed between HL and VI specimens. However, the number
of impacts sustained by the specimen before the onset of unstable growth is
smaller for VI specimens.
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ACKNOWLEDGMENTS
The authors wish to acknowledge AZIMUT Yachts for supplying the specimens.
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REFERENCES
[1] Roy, R., Sarkar, B.H. and Bose, N.R., 2001, “Impact fatigue of glass fibre-
vinylester resin composites”, Composites: Part A, 32, pp. 871-876.
[2] Azouaoui, K., Recha, S., Azari, Z., Benmedakhene, S., Laksimi, A. and
Pluvinage, G., 2001, “Modelling of damage and failure of glass/epoxy composite plates
subject to impact fatigue”. Int J Fatigue, 23, pp. 877-885.
[3] Wyrick, D.A. and Adams, D.F., 1998, “Residual strength of a carbon/epoxy
composite material subjected to repeated impact”. J. Composite Materials, 22, pp. 749-
765.
[4] Baucom, J.N. and Zikry, M.A., 2005, “Low-velocity impact damage progression
in woven e-glass composite systems”. Composites: Part A, 36, pp. 658-664.
[5] Kawaguchi, T., Nishimura, H., Ito, K., Sorimachi, H., Kuriyama, T. and
Narisawa, I., 2004, “Impact fatigue properties of glass fiber-reinforced thermoplastics”.
Composite Science and Technology, 64, pp. 1057-1067.
[6] Liu, D., 2004, “Characterization of impact properties and damage process of
glass/epoxy composite laminates”. J. Composite Materials, 38, pp. 1425-1442.
[7] Belingardi, G., Cavatorta, M.P. and Paolino, D.S., “A new damage index to
monitor the range of the penetration process in thick laminates. Submitted to
Composites Science and Technology.
[8] Belingardi, G., Grasso, F. and Vadori, R., 1998, "Energy absorption and damage
degree in impact testing of composite materials", Proceedings XI ICEM (Int. Conf.
Experimental Mechanics), Oxford (UK), pp. 279-285.
[9] Belingardi, G. and Vadori, R., 2003, “Influence of the laminate thickness in low
velocity impact behaviour of composite material plate”. Composite Structures, 61, pp.
27-38.
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[10] Belingardi, G., Cavatorta, M.P. and Paolino, D.S., 2006, “Repeated impact
behaviour and damage progression of glass reinforced plastics”. 16th European
Conference of Fracture (ECF16), Alexandroupolis, Greece, July 3-7, 2006.
[11] ASTM D3029 – “Standard Test Method for Impact Resistance of Rigid Plastic
Sheeting or Parts by means of a Tup (Falling Weight)”. American Society for Testing
Materials (1982).
[12] Schoeppner, G.A. and Abrate, S., 2000, “Delamination threshold loads for low
velocity impact on composite laminate”. Composites: Part A, 31(9), pp. 903-915.
[13] Choi, H.Y., Wu, H.Y.T. and Chang, F.K., 1991, “A new approach towards
understanding damage mechanisms and mechanics of laminated composites due to low-
velocity impact: Part II – analysis”. J. Composites Materials, 25(8), pp. 1012-1038.
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Table 1. Stacking sequence of the laminate. Thickness (t) of each ply is the one given in
the design protocol.
ply type t (mm) N. of plies
mat 0.74 2
UD 1.25 2
BD 0/90 1.40 2vinylester
mat 0.74 1
BD 0/90 1.40 2
UD 1.25 1polyester
mat 0.74 1
total t (mm) 12.31
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FIGURE CAPTIONS
Figure 1. An example of force vs. time and energy vs. time plots. Definition of force
and energy terms.
Figure 2: An example of quasi-static penetration tests and impact tests to perforation.Definition of the stroke variables sMAX and sQS.
Figure 3. Splitting of laminae at the interface between the two resin systems. Impacted
face: vinylester.
Figure 4. Testing fixture for impact testing
Figure 5. Representative force-displacement curves for quasi-static perforation tests.
Complete curve (a); initial portion for calculation of the laminate stiffness (b)
Figure 6. Representative force-displacement curves for repeated impact tests. Impact
velocity: 3.1 m/s. HL (a); VI (b)
Figure 7. Laminate stiffness against number of impacts.
Figure 8: Average laminate stiffness (a) and delamination Threshold Load –DTL- (b) as
a function of impact velocity.
Figure 9. Comparison of the peak force (a) and DI (b) vs. impact number. No
perforation tests.
Figure 10. Comparison of the peak force (a) and DI (b) vs. impact number. Perforation
tests.
Figure 11. Values of total impact energy density at perforation for different impact
velocities.
Figure 12. Slope of the DI vs. impact number curves plotted against the impact energy.
Figure 13. Pictures of a HL specimen impacted at 3.1m/s.
Figure 14. Pictures of a VI specimen impacted at 3.1m/s.
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Figure 1. An example of force vs. time and energy vs. time plots.Definition of force and energy terms.
0
5000
10000
15000
20000
25000
30000
0.0 1.0 2.0 3.0 4.0 5.0 6.0 7.0 8.0 9.0 10.0
Time [ms]
F o r c e
[ N ]
0
20
40
60
80
100
E n e r g y
[ J ]
Force
Energy
Impact
Energy
Peak
Force
Absorbed
Energy
Rebound
Energy
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Figure 2. An example of quasi-static penetration tests and impact tests to perforation.
Definition of the stroke variables sMAX and sQS.
0
5
10
15
20
25
30
0 8 16 24 32
Displacement [mm]
F o r c e
[ k N ]
0.2mm/s
3.8m/s
0
1
2
3
4
5
6
25 26 27 28 29 30 31 32 33 34
F~constant
sMAX=sqs
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Figure 3. Splitting of laminae at the interface between the two resin systems.
Impacted face: vinylester.
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φ 20
76.2
Stroke
A
B
CLAMPING AREA
Z
Y
X
Figure 4. Testing fixture for impact testing
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(a)
(b)
Figure 5. Representative force-displacement curves for quasi-static perforation tests.
Complete curve (a); initial portion for calculation of the laminate stiffness (b)
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(a)
(b)
Figure 6. Representative force-displacement curves for repeated impact tests.
Impact velocity: 3.1 m/s. HL (a); VI (b)
0
5000
10000
15000
20000
25000
30000
0 0.002 0.004 0.006 0.008 0.01 0.012
Displacement [m]
F o r c e
[ N ]
DTL= 11.6 kN
0
5000
10000
15000
20000
25000
30000
0 0.002 0.004 0.006 0.008 0.01 0.012
Displacement [m]
F o r c e
[ N ]
DTL= 12.8 kN
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Figure 7. Laminate stiffness against number of impacts.
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(a)
(b)
Figure 8: Average laminate stiffness (a) and Delamination Threshold Load –DTL- (b)
as a function of impact velocity.
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(a)
(b)
Figure 9. Comparison of the peak force (a) and DI (b) vs. impact number.
No perforation tests.
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(a)
(b)
Figure 10. Comparison of the peak force (a) and DI (b) vs. impact number.
Perforation tests.
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Figure 11. Values of total impact energy density at perforation for different impact
velocities.
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Figure 12. Slope of the DI vs. impact number curves plotted against the impact energy.
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1st
impactimpacted face
1st
impactrear face
impact before perforationimpacted face impact before perforationrear face
Figure 13. Pictures of a HL specimen impacted at 3.1m/s.
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impactimpacted face
1st
impactrear face
impact before perforationimpacted face impact before perforationrear face
Figure 14. Pictures of a VI specimen impacted at 3.1m/s.