IMPACT DAMAGE TO TEXTILE COMPOSITES

G. Zumpano, M. Sutcliffe, W.J. Stronge, M. Meo

CUED, Cambridge University, Trumpington steet, CB2 1PZ

Abstract

Textile composites offer an enhanced damage tolerance to impact when compared to conventional composite laminates, since their fibres are also oriented out-of-plane according to repetitive patterns defining the textile Unit Cell (UC). Objective of this study was to set up a benchmark for future studies on 3D weave composites by analysing the damage evolution on 2D woven composites. At this aim, high strength carbon fibre plane weave specimens of two sizes (20x20 and 13x13cm) were impacted by a drop weight with kinetic energy ranging from 4 to 18 Joules. Post-impact damage of target plates was assessed using thermography and sectioning. The analysis of the experimental results brought to conclude that the impact location with respect to the UC geometry is capable of introducing large changes in terms of damage severity. Finally, two main failure modes were identified: kinking failure on the impacted surface and tow bending/tensile failure combined with a shear failure of the matrix on the distal surface.

1. Introduction

Textile composites recently have been subject to increasing interest by the aerospace industry because of their enhanced damage tolerance to impact when compared against composite laminates. This increased damage tolerance to impact is related to out-of-plane fibres occurring in a repetitive pattern that defines the textile Unit Cell (UC). So far, research on textile composites has been focused on studying the parameters affecting impact damage such as projectile properties (impactor mass, shape and speed), sample dimensions, clamping conditions, UC geometry and materials.

In particular Robinson and Davies [1] analysed the effect of impactor mass at low velocity impact claiming that “the changes in the dynamics of the specimen/impactor system, due to changes in the impactor mass, do not significantly affect the stress distribution at the impact point”. Furthermore, the peak force was also shown to correlate with impact energy independently from the impactor mass, indicating that the low velocity impact of these specimens is a quasi-static process. A similar dependency of the damage progression with the impact kinetic energy was obtained by Gao and Kim [2], who studied the effect of fast cooling on a carbon reinforced thermo-plastic resin (PEEK) under impacting loading. The mass was kept constant, while the speed was changed. As a result the damage area increased with the impact energy closely resembling the behaviour observed by Robinson and Davies [1].

On the other hand, Naik et al. [3], investigating the effects of mass and speed by keeping constant the kinetic energy, observed that the impactor having high mass and low velocity combination yielded to the largest damage area, in contrast with the impactor having low mass and high velocity combination, which caused the smallest damage.

Projectile shape influence on damage extension due to low speed impacts was studied by Shim and Yang [4]. They observed that a sharp impactor produced more damage than a blunt impactor. Gellert et al. [5] highlighted a bilinear behaviour of the kinetic energy at the ballistic limit against the thickness and the projectile shape. A similar behaviour was observed by Mateminola and Stronge [6] on woven composite strips. Experiment results highlighted that the kinetic energy density for damage initiation had a maximum for impactor radius (r) and strip thickness ratio (h) equal 1 and that damage initiated on either impact side or the distal side of the strip according to r/h being smaller than or larger than 1.

Gellert et al. [5] also observed that penetration damage through thickness of plates having slenderness ratio (plate width – w – over plate thickness - h) less than 11 had a hourglass shape (convergent divergent shape), while for plates with a slenderness ratio larger than 11 the damage shape was observed diverging starting from the impact surface.

The effect of different woven composite geometries (plane and satin woven, stitched and unstitched) around the ballistic limit were investigated by Hosur et al. [8], who highlighted that the impact location could affect the damage evolution during the impact phenomenon. In particular, if the projectile impacted on stitch lines, the composite sheared easily, because of local damages caused by the stitching needle. Moreover, the satin weave had better ballistic speed than the plane weave, due to its geometry. Furthermore, it was observed that the stitching stopped splitting damage propagating, this determined smaller damage and, therefore, a reduction of the ballistic limit.

X. S. Zeng et al. [9] observed that absorbed energy and so the ballistic limit changed with different boundary conditions. In particular, for striking velocities below 240 m/s, fabric targets clamped along only two edges, generally exhibited superior energy dissipation compared to fabric clamped along four edges. In this regime, a fabric target clamped along two edges was capable of absorbing 90% more energy on average than one clamped along all four edges. In the high-speed penetration regime, the energy absorbed by targets clamped along two edges dropped significantly and fabrics clamped at four edges showed better performances.

Liu et al. [10] investigated the sample size effect observing that the most important parameter was given by plate thickness. Specifically, they observed a nonlinear dependence on the thickness which was thought to be related to the plate rigidity. Moreover, despite the slenderness ratios of the analysed plates were almost identical, the experiments showed very different behaviour between large and small samples.

All these information were taken into consideration for the setup of low speed impact experiments, which results were discussed in this paper. In particular, the impact kinetic energy was taken as main parameter of the investigation, while other factors as impactor weight and shape, plate thickness, etc… were kept constant. Moreover, the

sample sizes were chosen in order to have slenderness ratios lager than 20, in order to avoid occurrence of different mode failures as observed by Gellert et al. [5].

In this paper, the preliminary findings of a wide project aimed at investigate the low/high speed impact behaviour of 3D textile are presented. Specifically, the 2D plane weave textile composite low speed impact behaviour was investigated in order to establish modus operandi and comparison data for the experiments that will be carried out on 3D weave composites.

2. Material manufacturing

In order to reduce the scatter due to manufacturing variability, all the samples used in this investigation were cut out from a 100x50x0.2cm plate. The plate was manufactured using 10 layers of Hexcel prepreg plane weave sheet (HexPly® M47N/42%192P/CHS 3K) pre-impregnated with M47 resign (nominal weight 330g/m2, density 1,24 g/cm3) reinforced with high strength carbon fibres (nominal weight 192 g/m2, fibre density 1.78g/cm3). The fibres were organised in bundles of 3K, named tows (1.6mm wide), according to a repetitive pattern, plane weave (4x4x0.2 mm, Figure 1).

Figure 1 – Plane weave geometry.

The plate was cured in the CUED autoclave with a cycle of 90 minutes at 140 C and 6 bar pressure.

After curing, 2 5×13 (strips) and 6 13×13cm (small plates), 2 20×20cm (large plates) samples were cut out from the cured plate using a diamond saw. With the aim of verifying the UC geometry after curing, pictures of the plates were taken and processed using image recognition software based on an image intensity algorithm implemented in MATLAB® [Alex]. This software confirmed the estimate, from the pre-preg samples, for the UC length (4 mm) and the tow width (1.6mm, see Figure 2).

Figure 2 – Photo of the specimen used for UC length and tow width measurements.

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Moreover, the tow width estimate was also confirmed (at least, on average 1.55 mm) by the image analysis of the plate cross sections that were previously polished. Unfortunately, because of the polishing low quality (Figure 3), large data scattering was predicted for the tow length (standard deviation 11%, see Figure 4), while the tow area was predicted to be 0.145 mm2, with just 1% standard deviation. This means that the fibre volume ratio for the tow is 78.9% against the 58% of the pre-impregnated sheet.

However, this data must be considered a coarse estimate, since with better polishing and larger image magnifications (Figure 5), the tow shapes, fibre and resin pocket locations can be clearly identified and the image recognition software much more accurate.

Figure 3 – Cross-.

0.12

0.13

0.14

0.15

0.16

0.17

0.18

1 1.2 1.4 1.6 1.8 2

Tow width [mm]

Tow

are

a [m

mT

ow A

rea

[mm

2]

Figure 4 - Measured Tow area against Tow length.

Figure 5 - Particular of fibre locations in tow sections.

Finally, in terms of UC geometry, it can be seen from cross section pictures (Figure 3, Figure 5) that several irregularities are introduced by the manufacturing process such

1mm

200μm

as tow nesting, tow waving, deviation from lenticolar shape theorised by Kuhn and Charalambides [11] and etc…. These irregularities may play an important role in the impact behaviour of the textile composite, conditioning the damage evolution and, therefore, introducing a further source of scatter for the experimental data.

2. Low speed impact experiment procedure

Low speed impact experiments were carried out using Cambridge drop tower fitted with a drop weight of 0.62 kg and a hemispherical head of 12.5 mm diameter (Figure 6). The impacted samples were 10 plain weave composite plates: two 200×200 mm plate (PWs1L – Plane Weave Sample 1 Large –, PWs2L), six 130×130 mm plates (PWs3S – Plane Weave Sample 3 Small –, PWs4S, PWs5S, PWs6S, PWs7S and PWs8S) and two 50×130 mm strips (PWs6RSt – Plane Weave Sample 6 Right Strip –, and PWs6LSt were cut out from the post impacted PWs6S).

Two square plate sizes were investigated in order to take into account size effects on the impact dynamics, while the strip samples were investigated to highlight the effect of the boundary conditions and slenderness ratio. These specimens were clamped along their shortest edges, in contrast with the squared plates, which were clamped along all their sides.

The drop height was adjusted in order to have impact energies ranging from 4J (3.6m sec-1) to 18J (7.6m sec-1).

Figure 6 - High speed camera picture frame.

The procedure followed for low speed impact experiments involved 6 phases: clamping, pre-impact profile analysis, impact, post-impact profile analysis, NDT analysis, and sectioning.

The clamping phase involved the use of a clamping device (Figure 7), whose bolts were tightened to a torque of 20Nm. Afterwards, the clamped specimen profiles (upper and lower surfaces) were recorded by using an 3D co-ordinate measuring machine (Omicron: 40-50 µm resolution). Reasons for this analysis were to measure any curvature induced by clamping and to compare surfaces before and after impact in order to identify any sideways slip of the plate due to the impact.

Impactor head

Force Transducer

The impact dynamics was recorded using high speed camera, force transducer and PZT patch sensors. The impact and recoil speeds of the impactor were measured using a high speed camera (Figure 6). This allowed the estimate of the absorbed energy by the sample during impact. On the other hand, the contact force was monitored by a force transducer placed in contact with the impactor head (Figure 6). Furthermore, PZT patch sensors were positioned on the specimens in order to monitor the stress propagation during the impact event.

Figure 7 - Clamping device.

After impact, a second surface scan was carried out using the Omicron co-ordinate measurer.

The NDT phase consisted in thermography inspections of the impacted samples. Finally, the samples were reduced to 3×5 cm coupons, which were sliced at intervals of 1.2mm using an Accutom50 microtome. Afterwards, the slices were analysed under an optical microscope in order to identify impact damages, in terms of their location extension and type.

3. Sample pre and post profile analysis

The acquisition of the sample upper and lower surface profiles was carried out to have an estimate of an eventual curvature induced by the clamping on the sample, and to have a first non destructive estimate of the impact damage by comparing the pre and post impacted surfaces.

The upper and the lower sample surfaces were measured point by point on an irregular grid by the Omicron. Then, the 3D coordinate points were interpolated on a regular grid in order to plot them as whole on a surface (Figure 8). As result of this process, it was clear that the clamping induced a curvature on the samples with a maximum arc sagitta of 0.5mm. This meant that the curvature radius had a minimum of 5.6m and, hence, it was considered to be negligible.

Figure 8 – Pre-impact deformation of plate form clamp: upper and lower surfaces.

By analysing the post-impact profiles, an estimated of the visible damage is carried out (Figure 9), in terms of its extension, depth on both the plate surfaces: impact and distal. Moreover, it possible to assess if any side slip (Figure 9b and c) has occurred during impact by comparing the profiles before (dashed lines) and after impact (continuous lines).

Figure 9 – Post-impact plate deformation (PWs5S – 11.5J): (a) 3D representation; (b) plate cross section // to Y axis at the maximum deformation; (c) plate cross section // to X axis at the maximum

deformation.

4. Impact monitoring

The plate behaviour during impact was monitored through a series of devices aimed at quantifying the impact energy and sample absorbed energy (high speed camera),

Damage

(a)

(c)

(b)

measuring the contact force dynamics (force transducer) and impact stress wave propagation (PZT patch sensor).

The impact energy and sample absorbed energy were estimated from the impactor speed during impact. This was carried out by processing high speed camera picture frames (256×256 pixels) using an in house made program. Hence, by evaluating the speed immediately before impact (V0) and immediately after (Vf), the Impact Energy (IE) and the sample Absorbed Energy (AE) were evaluated as follows:

( )

20

2 20

1212 f

IE mV

AE m V V

=

= − (1)

with m the mass of the impactor (0.62 kg).

5. Thermography

Thermography images of the impacted samples were acquired in collaboration with Bath University using an EchoTherm system fitted with an infrared camera. The thermography imaging process used in this case is pulse thermography. This consists in the generation of high intensity light pulse that heats up the sample surface, which temperature is monitored by an infrared camera. The damage locations will have different thermal conductivity coefficients from the undamaged ones, appearing at first warmer than the surrounding undamaged locations, providing in this way 2D maps of the impact damage distribution. For each sample, two Thermography Images (TI) were taken (upper and lower surface), since the technique is insensitive to damages deeper than 2mm (Figure 10). Then, the TI were post-processed in Matlab in order to introduce a scale and to measure the damage area.

(a) (b)

Figure 10 - TI of PWs3S (7.5J): a) lower surface; b) upper surface.

By observing the TI (Figure 10), the damage on both surfaces was clearly identified, in term of its 2D extension. The damage on the bottom (Figure 10a) surface was always larger than the damage on the top surface and showed a very irregular perimeter. For low impact energy (up to 8J), damage on the top surface was mostly concentrated in few locations (dark spots in Figure 10b) that, as it will be clear from the micrograph images, coincided with surface or close-to-surface cracks. Moreover, the direction of tows on the two surface layers can be observed.

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6. Discussion

By analysing and comparing the data gathered by the previously illustrated procedures, useful remarks were made.

As expected boundary conditions affected the plane weave behaviour [9]. For the strip samples, the Absorbed Energy (AE) against Impact energy (IE - Figure 12), upper bounded the small plate AE, mainly, due to sample side slip (Figure 11), larger damages (Figure 13) and deflections induced by the damage compared to full clamped plates. In particular, strips tended to absorb on average 100% more than small plates in line with the 90% observed by Zeng et al. [9].

For small plates (PWS) in Figure 12, the AE ranges from 20% of the IE to 100% (PWs7S). In this case, the friction forces between the sample and the impactor head were such to halt the impactor run and require a considerable amount of force to untangle the two.

In terms of damage severity, small plates resulted upper bounded by strips and lower bounded by large plates, confirming the sample size effect observed by D. Liu et al. [10].

By observing the behaviour of the AE (Figure 12) and of the Contact Force Maximum (CFM - Figure 14), it is possible to remark that the AE drop associated to sample PWs4S cannot be due to normal scatter, since the sample CFM was in line with the behaviour of the remaining investigated samples. The same incongruence is also highlighted by Figure 13, where the damage area behaviour with the impact energy has been plotted. This means that the AE drop has to be associated with a smaller damage. The causes for such behaviour can be many. Though, looking at the impacted regions of the samples, a major player to this phenomenon seemed to be the impact location. Close ups of impact surfaces (Figure 15a) showed that, mostly, the impact centre interested mainly a single tow, as showed for sample PWs5S in Figure 15 (black marker line). However, in the case of sample PWs4S, this did not happen and the impact centre was located between two tows (Figure 15b).

Figure 11 – PWs6RSt - upper surface omicron profile: 1) pink surface undamaged profile; 2) coloured surface damaged profile.

0

20

40

60

80

100

0 3 6 9 12 15 18Impact energy [Joules]

PWs2LPWs3SPWs4SPWs5SPWs6SPWs6LStPWs6RStPWs7SPWs8SPWSA

bsor

bed

ener

gy [

IE%

]

Figure 12 – Absorbed energy vs impact energy.

0

50

100

150

200

250

300

350

0 3 6 9 12 15 18Impact energy [Joules]

PWs1LPWs2LPWs3SPWs4SPWs5SPWs6SPWs6RStPWs7SPWs8SPWSD

amag

e ar

ea [m

m̂2]

Figure 13 - Damage area evolution with the impact energy.

Damage initiation Plate perforation

Plate perforation

0.0

0.8

1.6

2.4

3.2

4.0

0 3 6 9 12 15 18Impact energy [Joules]

PWs2LPWs3SPWs4SPWs5SPWs6SPWs6LStPWs6RStPWs7SPWS

Max

imum

con

tact

forc

e [k

N]

Figure 14 – Contact force peak vs. impact energy.

(a) (b)

Figure 15 – Impact surface close ups: (a) PWs5S 11.5J; (b) PWs4S 12.5J.

Concluding, the results highlighted in this investigation confirmed literature data regarding to size and boundary effects, but also underlined the large influence of the impact location relative to the UC geometry on the damage severity. Therefore, this aspect will be the main focus for further investigations that will be carried out on 3D weaves, where the nature of UC might lead to the identification of different failure modes according to the impact location.

7. Sectioning

Sample sectioning provides useful information on the failure modes in such detail that none of the available today NDT techniques are capable of providing.

For low impact energy (4J-8J), moving from the boundaries of damaged areas towards their centre (Figure 16), kinking failures (Figure 17-a) of tows close to the top surface were identified as well as intra-tow cracks (Figure 17-b) between tows located close to the bottom surface. Approaching the centre of the damaged area, kinking failures vanish (Figure 18), though, at the same location, inter-tow cracks propagate through tow sections to deviate along tow surfaces (intra-tow crack). At the same time,

Impact centre

Plate perforation

the intra-tow cracks located at the bottom of the section enlarge encapsulating tow splitting cracks (Figure 19-Figure 20).

Figure 16 – PWs2L (5J): impact surface micrograph: ↔ fibre direction; − − − tow boundaries.

Moving away from the damage centre, kinking failures reappear on the impacted surface (Figure 21), while the damage at the distal surface decreases to vanish completely approaching the damage area boundaries.

By comparing the clear cut of the tow splitting crack due to kinking (Figure 17-a and Figure 19, upper surfaces) and the jagging cut of the broken fibres in the bottom inter-tow crack (Figure 20), it is clear that a different failure mechanism such as bending/tensile combined to a shear failure of the matrix (clear cut with an angle of about 45º with respect to the vertical) has occurred.

(a) (b)

Figure 17 – PWs2L (5J): a) kinking failure; b) intra-tow cracks.

Kinking failures Dent Area

(delamination)

Figure 18 – PWs2L (5J). Damage area centre: upper surface.

Figure 19 – PWs2L (5J). Damage area centre: lower surface

Figure 20 - PWs2L (5J): Damage area boundaries – Lower surface.

Figure 21 – PWs2L (5J): Damage area boundaries – Upper surface.

0.5mm

By using close-up pictures and thermal images, insights on the damage evolution with the increase of the impact energy can be drawn. On the upper surface (Figure 22), damage starts evolving from the impacted area in the shape of localised tow failures and matrix cracks along tow surfaces. The number of cracks of this fracture system increases with the impact energy and tend to be located at the boundaries of the indentation area (4-8J). A further increase of impact energy (11-18J) causes the surfacing of the inner cracks close to the impacted surfaces (Figure 18) leading to perforation, when these cracks and the lower surface cracks meet at the middle of the plate cross section. On the distal surfaces, failure starts in correspondence of the indentation area centre in the form of intra-tow cracks leading to tow splitting cracks with the increase of the impact energy.

Figure 22 – Damage evolution: Upper surface.

Impact surface

Distal surface

4.2J 7.5J 8J

4.2J 7.5J 8J

11.5 17.2J

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Figure 23 – Damage evolution: Distal surface.

Hence, damage was observed initiating on both impact and distal surface in contrast with S. A. Matemilola, W. J. Stronge [6] that observed the damage initiating either on the impact surface or the distal surface. Moreover, the damage on the impact surface was characterised by kinking and intra-tow failures, while on the distal surface intra-tow cracks (delamination) are predominant together to tow splitting. Finally, for low level impact energy (4-8J), the damage on the impact surface is concentrated in tows and, therefore, in some way affected by the impact location within UC geometry. On the other hand, on the distal surface, the damage is dominated by delamination.

8. Conclusions

In this paper, the results of the behaviour at low speed impact of high strength carbon fibre plane weave textile composites were investigated as preliminary case study before a comprehensive investigation on 3D weaves is carried out.

The steps followed during the impact experiments were described and the results reported giving a wide picture on the evolution of damage with the increase of impact energy for plane weave textiles.

The importance of the cross analysis of the impact monitoring data with the thermography results was highlighted by the fall in the absorbed energy and damage area of sample PWs4S with respect to sample PWs5S, nonetheless, the impact energy and contact force maximum were rising. This incongruence brought to discard the idea of data scatter and led to a closer analysis of the samples, concluding that the impact location was the discriminating factor. However, this insight needs further corroboration by an ad hoc experimental campaign.

Finally, the damage evolution with the impact energy was reconstructed by the combined analysis of thermography images, sectioning and surface micrographs. This stressed that there were two different main failure modes: (a) tow splitting and (b) bending/tensile failure combined to a shear failure of the matrix, occurring respectively on the impact surface and on the distal surface.

11.5 17.2J

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12.