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-structured sandwich composites response to low-velocity impact Antonio F.
, 
, , Maria Gabriela R.
This paper investigates the influence of exfoliated 
nano-structures on sandwich composites under impact loadings. A set of sandwich composites plates made of fiberglass/nano-modified
epoxy face sheets and polystyrene foams was prepared. The core was
25 mm thick and the face sheets were made of eight layers of woven
fabric glass fibers and nano-modified epoxy (≈0.8 mm of thickness). The epoxy system was bisphenol A resin and an amine hardener. The fiber
volume fraction used was around 65%, while the nanoclay content varied
from 0 wt.% to 10 wt.%. The nanoclay used was Cloisite 30B from
Southern Clay. The sandwich panels were submitted to low-velocity
impact tests with energies from 5 J to 75 J. Two sets of experiments
were performed, i.e. high velocity + low mass and low velocity + high
mass. Damage caused by the two groups of experiments and peak forces
measured were dissimilar. The results show that the addition of 5 wt.%
of nanoclay lead to a more efficient energy absorption. The failure
modes were also analyzed, and they seems to be affected by the nanoclay
addition to face sheets.
Keywords: Sandwich composites; Nanocomposites; Low-velocity impact; Nanoclay; Mechanical properties; Modeling
As said by Mines and his co-workers [5], the low-velocity impact of composite sandwich panels (to this investigation structures and panels have the same meaning) covers the response from sub-critical damage to full perforation. Furthermore, an important point of interest investigated by Mines et al. [5] was the relationship between the peak impact force and the damaged area. They concluded that core density has direct influence on damage propagation and energy absorption. Another issue, i.e. the sandwich composite rate-sensitivity, was investigated by Hazizan and Cantwell [6]. They demonstrated that foam based sandwich composites are insensitive to crosshead displacement rate over a large range (0.1–1000 mm/min). Ávila [7] has shown that by varying the sandwich composite core density, as in functionally graded sandwich structures, damage mechanism also changes as function of density distribution. Laurin and Vizzini [8] went further, as they demonstrated that by placing graphite/epoxy reinforcement through-thickness in a foam core enhances the normal energy absorption of foam core sandwich structures. Still, they called our attention to the fact that there is no linear relationship between the increase on energy absorption and the addition of stiffness introduced. Suvorov and Dvorak [9] also tried to enhance the sandwich energy absorption by placing an interlayer material between the face sheet and the foam core. Their mathematical model seems to be encouraging. Nevertheless, no experimental data was provided to validate the model.
Ávila et al. [10] showed that failure mechanisms of laminated composites (in the present investigation face sheets) can be influenced by nano-structures formed by nanoparticles dispersed into epoxy systems. According to them, the presence of nanoclay into fiber
glass/epoxy composites lead to a more intense formation of delaminated
areas after a low-velocity impact test (LVI). This phenomenon was
attributed to interlaminar shear forces caused by the intercalated nano-structures
inside the epoxy system. Furthermore, the energy absorption of these
laminates increased by 48% with dispersion of 5 wt.% of nanoclays.
Another important issue described by Li et al. [11]
is that, when sandwich composites were submitted to LVI tests, there
was impact weave propagation inside the two distinct materials, i.e.
face sheets and core. Ávila et al. [12] demonstrated that nano-modified laminated composites had their natural frequencies altered by nano-structures formed inside the laminated. A crossing mode, 4th mode higher than the 3rd mode, was reported by Ávila et al. [12] with the addition of 5 wt.% of montmorillonite to fiber glass/epoxy laminates.
This paper focus on the effect of nanoclay dispersed into an epoxy/fiber glass composite and its influence on the sandwich composites response to low-velocity impact tests.
Ten sandwich panels were evaluated during the course of this research. The face sheets of the sandwich panel were fiber glass/nano-modified epoxy system laminates. The epoxy system was made of bisphenol A
resin and an amine hardener, i.e. RemLam M and HY956, from Hunstman
Inc. The nanoclay used in this research was organic modified
montmorillonite produced by Southern Clay Inc. (Cloisite 30B). The
nanoclay content employed was 0 wt.% (control), 1 wt.%, 2 wt.%, 5 wt.%
and 10 wt.%. Eight layers of plain weave woven fabric fiber glass with aerial density of 200 g/m2 were employed, resulting in a cured plate with thickness of 0.8 mm and a fiber volume ratio of 0.65. The core used for the sandwich panel was closed-cell PS foam (Styrofoam®). This foam has a density of 35 kg/m3.
The core was 25.4 mm thick. During the face sheet preparation, the
nanoclay dispersion process into the epoxy system followed the
procedure describe in Ávila et al. [13].
A dispersant agent, acetone, was employed to improve the mixing
process. The degassing stage was required to eliminate bubbles
generated during the mechanical mixing, as well as to eliminate the
dispersant agent, i.e. acetone. After this procedure, the hand lay-up
with vacuum assisted cure was performed. The sandwich panel was
fabricated by bonding the cured face sheets to the core material with
the same epoxy system applied during the laminate consolidation. A
vacuum bagging technique was used to ensure a uniform pressure during
the bonding procedure. The final sandwich panel was a square plate of
300 × 300 mm dimensions with an overall thickness of 27.0 mm and mass
of 0.52 kg.
As described by Koo [14], during the nanoparticles dispersion into polymeric matrices nano-structures are formed. The two most common detection techniques to nano-structures
identification are X-ray diffraction and electronic microscopy. In this
research, X-ray diffraction (XRD) experiments were carried out on a
Shimadzu XRD-6000 X-ray diffract meter with Cu (λ = 0.154 nm)
irradiation at 40 kV and 30 mA using a Ni filter. Data were recorded in
the range from 2 to 80 deg in a continuous scanning at 2 deg per minute
and sampling pitch of 0.02 deg. The scanning electron microscope (SEM)
used was a LEO model 1430VP. In all cases, a gold thin film was placed
on the surface of the sample to able to scan its structure. Finally, a
thermal stability analysis was also performed using a Shimadzu DTG 60
Thermogravimeter under nitrogen, at 10 °C/min from 25 °C to 750 °C. As
this research deals with dynamic loadings, in addition to the
traditional nanoparticles dispersion analysis, information about the
sandwich response under indentation must be supplied.
According to Hazizan and Cantwell [6],
during impact events sandwich structures absorb significant energy in
contact deformations around the impact area. Thus, a series of
indentation tests were carried out using an INSTRON universal testing
machine at crosshead displacement rate of 0.2 mm/min. The indentation
tests were conducted using a hemi-spherical indentor of 10 mm diameter.
The sandwich panels were loaded up to a force with same magnitude of
the ones recorded during the low-velocity impact tests. As described by
Abrate [4], during the loading phase of the impact, the contact force P is related to the indentation α by
The load-indentation data were fitted to this equation to yield the values of n and C for each group of sandwich panels.
Once the indentation tests were performed, the following steps are the low-velocity impact tests and failure mode analysis. Following the ASTM D 5628-01 [15], the dart has a hemispherical nose with a radius of 10.0 ± 0.1 mm. The specimen is clamped between two rectangular steel plates with 13 mm thickness and a central circular cutout of 50.0 ± 0.1 mm. The dart has a weight of 246 g and six additional steel circular plates with a diameter of 75.0 ± 0.1 mm and a thickness of 15.0 ± 0.1 mm. They are placed into the rod linked to the dart. The average weight of each circular plate is 528 g. As the drop weight tower has a maximum height of 3.0 m, the limiting velocity for the device is 7.67 m/s. The dart is made of AISI 4330 steel. The six steel disks can be assembled individually into the dart leading to a mass variation ranging from 246 to 3414 g. During the impact event, the load–time history is recorded by a data acquisition system with a frequency of 50 kHz rate by an A/D data acquisition board. As stated by Found et al. [16], any digital filtering must be done at a cut-off frequency or below this value, which is in general half of the maximum data acquisition rate. However, as described by Ávila et al. [12], the cut-off frequency should be l0–15% of the bandwidth of the data in order to avoid problems. In the present case, after a series of pre-tests, a low-pass digital filtering was selected at 3.5 kHz, i.e. 14% of the maximum bandwidth.
In contrast with traditional low-velocity impact (LVI) tests performed, the dart is not embraced after the first impact. On the contrary, it is kept free so that it can strike back on the plate as many times as the plate reverses its direction of motion. By applying this strategy, it is possible to evaluate the amount of damage sustained by the sandwich panels. After the LVI tests the damage was measured using image correlation technique. Public domain image correlation software called ImageJ was employed. At least 10 measurements of each damage area were made and the average and standard deviation were calculated. To be able to predict the LVI peak force, an modification of the model described in Schubel et al. [1] is proposed.
According to Schubel et al. [1],
experimental data has confirmed that load pulse for LVI tests can be
expressed by a half-sine wave. Furthermore, for low and mid-range
energies the load pulse on sinusoidal shape can be expressed aswhere P0 is the impact peak load and T is twice the pulse duration. However, for high energies some perturbations must be introduced due to face sheet damage.
Now,
assuming the force–deflection for the LVI test having a quasi-linear
behavior and the peak force being represented by the superposition of
contact and bending forces, the following equation can be written aswhere the first term is the indentation component as described by Hazizan and Cantwell [6] and the second term represents the panel stiffness. Note that m is the impact mass and w the panel displacement under the tup.
To be able to identify the type of nano-structure
formed during the nanoclay dispersion process, a series of X-ray
diffraction (XRD) tests were performed. The data, shown in Fig. 1 indicate a high rate of exfoliation. The undulation curve shape showed in Fig. 1A is the contribution of the amorphous part of the fiber
glass micro-reinforcement used. Yet, the amount of nanoclay is twice as
much for the 10 wt.% nanoclay content specimens and the XRD intensity
increases only 13% (see Fig. 1B).
This unexpected reduction on the diffraction intensity is an indication
of a disordered swollen structure, as demonstrated by Fan et al. [17] Likewise, Ranade et al. [18]
stated that it could be an evidence of a small amount of nanoclay that
does not get exfoliated or intercalated, but remains in its own
identity leading to an immiscible nano system. This seems the case for the 10 wt.% concentration samples. This hypothesis can be visualized in Fig. 1C where the presence of these immiscible nano systems clusters (white regions) and an intercalated nano-structure can be observed. As stated by Luo and Daniel [19]
a partially exfoliated/intercalated system is most likely to occur due
to the dispersion system used. In summary, they believed that
dispersion process is controlled not only by the nanoparticle
morphology but it is also affected by the manufacturing process.
| Full-size image (81K) |
Cloisite nano-structures; (a) XRD signature; (b) XRD signature zoom; (c) cluster formation of precipitated non-reacted nanoclay.
The load-indentation depth curve for the face sheets were obtained based on the methodology described in Zenkert et al. [20] and the experimental load-central displacement curve. Following the same procedure applied by Tan and Sun [21], a nonlinear curve fitting is used to obtain the 1.5 power law equations. The results are shown in Fig. 2, while the curve fitting parameters are listed in Table 1. Furthermore, the C coefficients obtained are proportional to the square root of indenter tip, as predicted by Sutherland and Soares [22] and Shivakumar et al. [23]. Note that for sandwich composites, the n coefficient is lower than 1.5. Hazizan and Cantwell [6] attributed this lower value to the compliance of the test machine. A more reasonable explanation for the phenomenon was provided by Fatt and Park [24]. According to them, the sandwich panel response can be characterized by the superposition of two effects, that is, the bending of the upper face sheet and the core crushing in the areas below and around the indenter. The core crushing effect is the probable cause of this low value for the n coefficient.
| 30B wt.% | C [N/mmn] | n | Ra2 | Average residual |
|---|---|---|---|---|
| 0 | 286.6419 | 1.3397 | 0.9999 | 1.1085E−7 |
| 1 | 488.5669 | 1.2509 | 0.9998 | 1.7502E−7 |
| 2 | 573.2445 | 1.1771 | 0.9992 | 1.7766E−7 |
| 5 | 521.4524 | 1.1744 | 0.9991 | 1.8446E−7 |
| 10 | 413.9050 | 1.2690 | 0.9998 | 1.7937E−7 |
To be able to understand the impact response of sandwich composites with nano-modified
face sheets a series of LVI tests were performed. The energies employed
were enough to cause damage ranging from a barely visible delamination
and perforation. In all cases, the core material is the same; by
applying this strategy the core density effect is avoided. Table 2 and Table 3
show the LVI test parameters. Note that for each energy span, there are
two types of impact tests. The first one corresponds to a configuration
with the minimum mass (tup + load cell + support) and maximum velocity
(condition 1), while the second one is the opposite, i.e. maximum load
and minimum velocity. The objective is to investigate the impact
velocity and mass effects on damage formation and the sandwich panel
response to these variations (see Table 3).
| Energy (J) | Height (m) | Mass (kg) | Velocity (m/s) |
|---|---|---|---|
| 25 | 2.36 | 1.079 | 6.80 |
| 30 | 1.91 | 1.603 | 6.12 |
| 35 | 2.23 | 1.603 | 6.62 |
| 40 | 2.54 | 1.603 | 7.06 |
| 45 | 2.15 | 2.127 | 6.49 |
| 50 | 2.39 | 2.127 | 6.86 |
| 55 | 2.63 | 2.127 | 7.18 |
| 60 | 2.30 | 2.659 | 6.72 |
| 65 | 2.49 | 2.659 | 6.99 |
| 70 | 2.69 | 2.659 | 7.26 |
| Energy (J) | Height (m) | Mass (kg) | Velocity (m/s) |
|---|---|---|---|
| 25 | 0.69 | 3.717 | 3.56 |
| 30 | 0.82 | 3.717 | 4.02 |
| 35 | 0.96 | 3.717 | 4.34 |
| 40 | 1.09 | 3.717 | 4.64 |
| 45 | 1.23 | 3.717 | 4.92 |
| 50 | 1.37 | 3.717 | 5.19 |
| 55 | 1.51 | 3.717 | 5.44 |
| 60 | 1.65 | 3.717 | 5.68 |
| 65 | 1.78 | 3.717 | 5.91 |
| 70 | 1.92 | 3.717 | 6.14 |
| 75 | 2.06 | 3.717 | 6.35 |
Fig. 3A
gives the force versus time data for a 5 wt.% panel showing the effects
of impact velocity. For lower velocities, i.e. up to 4.34 m/s, the
impact period seems to be approximately the same. However, for higher
velocities (
5.19 m/s) a decrease on impact period was observed. As it can be noticed in Eq. (3),
stiffness is inversely proportional to the impact period. However,
taking into consideration the same panel stiffness, remember we are
dealing with panels with 5 wt.% nanoclay, the decrease on impact period
must be explained by the failure mechanism. Higher velocities imply
naturally higher energy and a much higher impulse (force/unit of time).
A higher impulse will cause a greater localized failure, mainly into
core. Core crushing will occur most likely below the tup. This
localized core crush contributes to the face sheet failure due to
absence of an elastic base. Velocities up to 4.64 m/s produced
sub-critical damage, i.e. before upper face sheet tearing. At low
velocities and consequently small energy levels, a large part of the
impact energy is easily dissipated by the face sheet first and the
remaining energy is absorbed by core. When the velocity increases to
around 5.19 m/s the upper skin tearing and core crushing is observed.
This failure mechanism can be explained by a large impulse which caused
a high localized damage, as there was not enough time to impact wave
propagation through regions outside below the tup area. This phenomenon
is more evident for high energy levels, i.e. >50 J, where the impact
wave propagates through the panel thickness right below the tup impact
and no significant damage was notice outside the impact projected area.
Upper skin tearing and lower skin damage was notice when the velocity
reached 5.44 m/s or higher. It is a well known fact that as the energy
gets bigger, the damage severity also increases. Perforation with lower
skin tearing was experimented for velocities higher than 6.99 m/s. Fig. 3B
allows us to investigate the mass effect on sandwich panel response to
low-velocity impact. The velocity was kept higher than 6.3 m/s while
the mass changed from 1.08 kg to 3.7 kg. The damage caused by these
impacts ranged from a barely visible (lower mass) to perforation. As
the velocity was kept near constant, the pulse duration could be
considered approximately the same. Another important issue is that face
sheet damage mechanism changes with the amount of nanoparticle
dispersed.
| Full-size image (42K) |
Force versus time for a 5 wt.% panel. (a) Impact velocity effect. (b) Impact mass effect.
Fig. 4A–C
provides the force versus time data for condition 2 where the energies
were 25, 50 and 75 J, respectively. A change on impact pulse is noticed
for the 25 J energy level. According to Eq. (5), impact pulse is
inversely proportional to panel stiffness. Thus, as the panel with
5 wt.% nanoclay content has the smallest period (26.89 ms), it must
have the largest stiffness. Another interesting issue is that pulse
period decreases with the addition of nanoclay up to 5 wt.%. The
10 wt.% panel has a pulse period 14.96% higher than the 5 wt.% panel.
The same pattern was noticed for the mid and high range energy, i.e.
50 J and 75 J, respectively. As described by Eq. (5), higher impact
period implies into smaller panel stiffness. Therefore, it is possible
to conclude that the “extra amount” of nanoparticle did not increase
the panel stiffness, in the contrary; it causes a decrease on
stiffness. One possible explanation for this phenomenon is the
existence of a “saturation” limit of the epoxy system. Once the
saturation limit is reached, the additional nanoclay dispersed into the
matrix precipitates into the form of immiscible nano-structures,
which can be sources of cracks nucleation, as they are stress
concentration “hot spots”. This hypothesis can be corroborated by the
XRD tests, where a significant increase on diffracted energy was
notice. Notice that a high diffracted energy is also a sign of large
entropy. This increase on entropy can be caused by these precipitated
third phase. This hypothesis was confirmed by micrographics using SEM (Fig. 1C).
| Full-size image (68K) |
Force versus time data. (a) Low range energy; (b) mid-range energy; (c) high range energy.
Damage observed on sandwich panels caused by the tup impact can be classified in six categories: (i) barely visible damage; (ii) sub-critical damage, i.e. upper face sheet cracking and core indentation; (iii) upper face sheet tearing and core crushing; (iv) lower face sheet and core debonding; (v) lower face sheet cracking; (vi) lower face sheet tearing and perforation. Fig. 5A–F shows each one of these categories.
The model predictions (Fig. 6) were compared against the experimental data with good agreement. However, for large energies (>50 J), a small deviation was observed. This can be due to the intense core crushing under the impact area. Notice that for the largest energy 75 J, a 6% difference between the predicted value and the experimental data is notice. The core crushing rate effect must be included in the present model.
The effect of nanoclay addition to fiber
glass/epoxy systems and its use on face sheet for sandwich composites
was investigated. X-ray diffraction tests and electronic microscopy
observation proven the existence of exfoliated nano-structures
dispersed into the epoxy system. The sudden drop on X-ray diffraction
intensity seems to be an evidence of entropy increase as result of an
immiscible phase precipitation. It seems that epoxy system has a
“saturation” limit between 5 wt.% and 10 wt.% of nanoclay content. A
model to predict the peak reaction for during the low-velocity impact
was proposed based on pulse duration. The model predictions are limited
as the core yielding was not taken into consideration. The addition of
nanoclay gives the impression of having effect on pulse duration, as
they increase the face sheet laminate stiffness and consequently the
panel overall bending rigidity. The pulse period appears to be
independent of velocity. Nonetheless, mass changes can affect the pulse
period. The nanoclay dispersion on epoxy system could be the reason for
an increase on face sheet stiffness and the panel overall rigidity when
low-velocity impact tests were performed. The nanoclay increase on
stiffness looks like to be overshadow by the core crushing during the
failure mode transition stages, i.e. from upper face sheet
cracking + core indentation to upper face sheet tearing + core
crushing, and upper face sheet tearing + core crushing to lower face
sheet and core debonding.
The first author would like express his gratitude to Dr. Claudia A. Vanetti and the Center of Microscopy and Microanalysis of Universidade Federal de Viçosa (NMM/UFV/FINEP) for the usage of the SEM. We also would like to acknowledge the financial support provided by the Brazilian Research Council (CNPq) under grants 470511/2006-0, 472213/2007-5 and 300434/2008-1.
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Corresponding author. Tel.: +55 31 3409 5238; fax: +55 31 3443 3783
