low velocity impact performance of recyclable all...

COMPOSITES

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Composites Science and Technology 66 (2006) 1724–1737

SCIENCE ANDTECHNOLOGY

Low velocity impact performance of recyclableall-polypropylene composites

B. Alcock a,*, N.O. Cabrera a,b, N.-M. Barkoula a, T. Peijs a,b

a Department of Materials, Queen Mary, University of London, Mile End Road, London E1 4NS, UKb Eindhoven Polymer Laboratories, Eindhoven University of Technology, P.O. Box 513, 5600MB, Eindhoven, Netherlands

Received 1 July 2005; received in revised form 5 November 2005; accepted 7 November 2005Available online 3 January 2006

Abstract

Highly oriented polypropylene (PP) tapes, with high tensile strength and stiffness achieved by molecular orientation during solid statedrawing are consolidated to create high performance recyclable ‘‘all-polypropylene’’ (all-PP) composites. These composites possess alarge temperature processing window (>30 �C) and a high volume fraction of highly oriented PP (>90%). This large processing windowis achieved by using co-extruded, highly drawn PP tapes. This paper investigates the impact resistance of these all-PP composites, and therelationship between penetrative and non-penetrative impact behaviour, and composite consolidation conditions. The response of all-PPcomposites to falling weight impact is reported together with a comparison to conventional commercial glass reinforced polypropylenecomposites. A model for energy absorption is proposed by comparison with previous studies based on interfacial and tensile failure oftapes and composites.� 2005 Elsevier Ltd. All rights reserved.

Keywords: B. Interface; B. Impact behaviour; A. Recycling; C. Delamination; A. Fabric/Textiles

1. Introduction

All engineering structures are susceptible to impactloading within their service lives. The response of the mate-rial to impact loading will depend on various factors suchas the geometry of the structure and striker, the massand velocity of the striker, and frequency of impacts.Due to their high strength and stiffness, and good energyabsorption due to delaminating failure modes, compositematerials generally perform well in impact applications.Carbon and glass fibres suffer from a lack of plasticitywhich means that non-penetrative impact loads can leadto (often invisible, subsurface) fibre damage, which candrastically reduce the residual mechanical properties ofthe composite [1]. Thermoplastic fibre composites typicallypossess sufficient elastic limits to make them less sensitiveto damage from lower energy impacts [2] (see Table 1).

0266-3538/$ - see front matter � 2005 Elsevier Ltd. All rights reserved.

doi:10.1016/j.compscitech.2005.11.010

* Corresponding author.E-mail address: [email protected] (B. Alcock).

Thermoplastic fibres such as UHMW-PE (e.g., Dyneema�,DSM, or Spectra�, Honeywell) have specific applicationsas impact defence materials, such as personal protectionfor military/police personnel from direct projectile impact[3], or as spall liners behind ceramic/metallic armour inarmoured vehicles to limit proliferation of shrapnel insidea vehicle following impact [4].

Composite ballistic protection can also provide signifi-cant weight savings for automotive defence, compared tosteel armour [5,6] and has also been assessed as fragmentbarriers for commercial aircraft [7]. The impact perfor-mance of composites has been modelled with some successto determine the methods of predict deformation [8] andmodel energy absorption [9].

1.1. All-polypropylene composite processing

In a series of PhD theses [10–12], novel composite materi-als in which both the fibre and the matrix are based on poly-propylene (PP) have been described. The creation of single

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Table 1Typical fibres properties for a range of commonly used fibre reinforcements for impact applications

Material Fibre type Tensile strength(MPa)

Tensile modulus(GPa)

Strain to failure(%)

Density(g cm�3)

Reference

PBO – 5.5 280 2.5 1.56 [3]

Glass E Glass 3.5 72 4.8 2.58 [50]S-2 Glass 4.9 87 5.7 2.46 [50]

Aramid Twaron HM1055 2.8 125 2.5 1.45 [51]Twaron HS2000 3.8 90 3.5 1.44 [51]Kevlar 49 2.9 135 2.8 1.45 [51]Kevlar 129 3.4 99 3.3 1.45 [51]

UHMW-PE (Gel processed) Dyneema SK60 2.7 89 3.5 0.97 [52]Dyneema SK71 4.0 120 4.1 0.97 [52]Spectra S900 2.1 79 3.6 0.97 [52]Spectra S2000 3 116 2.9 0.97 [52]

UHMW-PE (Melt processed) Certran 1.2 67 6 0.97 [53–55]

UHMW-PP (Gel processed) – 0.98 36 3.3 0.91 [56]

PP All-PP tapes 0.45 15 7.5 0.78 [11]

B. Alcock et al. / Composites Science and Technology 66 (2006) 1724–1737 1725

polymer composites is motivated by the desire to enhancerecyclability of composite materials. Conventional compos-ite employ very different materials for the matrix and rein-forcement phase and this complicates recycling. All-PPcomposites overcome this problem since at the end of the lifeof an all-PP product, the entirely polypropylene compositecan simply be melted down for reuse in a PP feedstock oreven in a subsequent generation of all-PP composite.

While neither the concept of creating high moduluspolymer fibres [13–23] or single polymer composites arenew [24–33], existing technologies have inherent limitationssuch as a relatively small temperature processing windowor a low volume fraction of reinforcement, which limitthe ultimate mechanical properties of the composites.Highly oriented, high modulus fibres or tapes can be effec-tively welded together by melting the surface of the tapesand applying pressure to achieve a good bonding and fillany voids [33–36]. In these mono-extruded tape or fibresystems, the process becomes highly sensitive to compac-tion temperature, since there is a risk of molecular relaxa-tion during the high temperature consolidation of tapes orfibre bundles into composites.

The research reported in this paper focuses on the use ofco-extruded tape technology to create all-PP compositeswith a large temperature processing window (>30 �C) andhigh volume fraction of reinforcement (>90%). The largetemperature processing window of these co-extruded tapesallows all-PP composite production over the range of thistemperature processing window by providing enough ther-mal energy to allow consolidation of tapes into a load bear-ing structure but without significant loss of mechanicalproperties. This does not imply that all mechanical proper-ties are uniform over this temperature window. The effectof compaction temperature and pressure on the mechanicalproperties of all-PP composites is discussed in great detailelsewhere [37,38]. These co-extruded tapes possess a skin-core morphology and are composed of a core of PP homo-

polymer sandwiched between a thin skin of a PP copoly-mer. This three-layer structure is co-extruded in a highviscosity melt phase and subsequently drawn in a two stagesolid state drawing process also described in greater detailelsewhere [11,22,39]. This drawing process results in a highdegree of molecular orientation and the drawn tapes pos-sess a high tensile strength (>450 MPa) and stiffness(>15 GPa). These co-extruded tapes can then be consoli-dated into a composite material by the application of heatand pressure to stacked plies of woven tape fabrics[12,37,39]. The application of pressure also causes a physi-cal constraining effect which has been shown to artificiallyraise the melting temperature of highly oriented polymers[22,40,41], and this effect further helps to retain the tensileproperties in the tapes by preventing molecular relaxationduring consolidation. The copolymer skin layer possessesa lower melting temperature than the homopolymer coreand hence allows tapes to be effectively welded togetherat temperatures far below the melting temperature of thehomopolymer core. The proportional thickness of the skinto the core can be altered during co-extrusion, but since theskin layer is present only to facilitate intertape bonding, itis desirable to have this skin as thin as possible whileachieving a high interfacial strength. The mechanical prop-erties of these tapes together with the inherent low densityof PP, and the high volume fraction of reinforcement pres-ent in these composites (Vf > 90%) make these all-PP com-posites competitive with conventional PP matrixcomposites [12,37,39]. Since the PP tapes used in thisresearch are co-extruded, the matrix phase (skin layer) iscarried by the core layer, in one tape. The optimisationof both the homopolymer (reinforcement) layer for highmechanical properties and the copolymer (matrix) layerfor good interfacial strength must be considered simulta-neously. While the extrusion and solid state drawingparameters determine the mechanical properties of thesetapes, the same parameters also affect the morphology of

1726 B. Alcock et al. / Composites Science and Technology 66 (2006) 1724–1737

the copolymer layer strength of the interfaces formed whenthese tapes are bonded to each other in a composite system.Energy absorption during impact is achieved by a varietyof failure modes, predominantly a combination of fibrefailure and interfacial failure. Since composite consolida-tion parameters control both the interfacial propertiesand the mechanical properties in woven tape all-PP com-posites, it is important to investigate how composite con-solidation parameters affect the impact performance ofall-PP composites [11].

This paper aims to investigate the impact energy absorp-tion by woven tape all-PP composites, processed with arange of processing conditions. By comparing the energyabsorbed during pure interfacial failure with energyabsorbed during pure tensile failure of tapes, a model isproposed which can be used to predict the impact perfor-mance of all-PP composites. The temperature and impactvelocity are kept constant in the experiments presented inthis paper, since the viscoelastic nature of polymeric sys-tems means that these factors are likely to effect impactenergy absorption.

1.2. Materials

To illustrate the versatility of all-PP composites, twotypes of co-extruded tapes will be used for composite pro-duction. Three-layer (skin:core:skin) tapes are co-extrudedand drawn in a solid state drawing process as described inTable 2. In this way, the material which will comprise thematrix phase is present in the tapes as a thin ‘skin’ on thetop and bottom of the reinforcement phase, the ‘core’ ofthe tape. Both of these tapes are composed of the samepolymers and only subjected to different drawing condi-tions. Both tapes consist of a PP homopolymer blend corelayer (Mw = �300 kg mol�1, MFI = 5 g min�1) co-extruded with a copolymer skin layer (Mw = 320 kg mol�1,MFI = 5.5 g min�1). Following co-extrusion, the tape isdrawn with parameters described in Table 2, to yield twokinds of tape designated A and B which are woven intofabrics A and B, respectively. During solid state drawing,molecular orientation and so high mechanical propertiesare achieved depending on draw ratio applied, and afterdrawing the tapes typically have a width of 2 mm and athickness of 65 lm.

All-PP fabric A is composed of a high draw ratio tape,with high tensile modulus, high strength and a relativelypoor interfacial strength. Fabric B is composed of a lowerdraw ratio tape with a lower tensile modulus and strengthbut a stronger interfacial strength. The interfacial strengths

Table 2Details of all-PP fabrics A and B

Fabric Tapedrawratio (k)

Tapemodulus(GPa)

Tapestrength(MPa)

Relative tapecomposition(skin:core:skin)

Drawingtemperature(�C)

A 17 15 450 5:90:5 170B 8 6 300 10:80:10 160

were characterised using single tape T-peel tests (ASTM1876) [11] and will be described in greater detail in a dedi-cated publication. These processing characteristics affectmechanical and interfacial properties and so greatly affectfailure modes. By using two different fabrics, the range ofproperties achievable with all-PP composites is demon-strated, as is the strong link between tape processing prop-erties and composite properties, allowing user-definabletape properties to control ultimate composite properties.

1.3. Specimen preparation

Penetrative falling weight impact testing employs lami-nates of material which are subjected to impact normalto the plane of the laminate as shown in Fig. 1. The pro-duction of all-PP laminates will now be described. Co-extruded all-PP tape is woven into a plain weave fabricwith an areal density of �100 g m�2. This fabric is cut intosquare plies and placed in close fitting mould with dimen-sions of 180 · 180 mm. This mould can then be placed ina hot press and heat (130–170 �C) and pressure (1–12 MPa) can be applied. Alternatively, this mould can beconsolidated in a vacuum bag while heated in an oven toachieve temperatures of 130–150 �C and pressure of 0.1–1 MPa. In either case, temperature in the mould is moni-tored using externally monitored PT100 temperatureprobes, and temperature across the mould was seen to beuniform within 1 �C; this is crucial since maintenance ofhigh mechanical properties depends on accurate tempera-ture control. The application of lateral pressure by eithera hot press or a vacuum oven acts to physically constrainthe fabrics and prevents relaxation during heating[22,40,41]. Upon cooling, pressure is removed and consoli-dation of the composite is complete. The all-PP compositecan then be removed from the mould. A range of consoli-dation temperatures and pressures was investigated todetermine the effect that these had on impact performance.

In addition to mechanical properties, composite densityalso varies with compaction parameters. With increasingcompaction pressure and temperature, superior consolida-tion of fabric plies is achieved and so the apparent densityof the composite laminates increases by the closure of inter-ply voids. For this reason, this paper often refers to specificmechanical properties, i.e., mechanical properties propor-tional to density. For example, a poorly compacted speci-men may appear to have a lower tensile modulus, but thismay be due to a higher volume fraction of interply voids giv-ing a falsely low density and hence an inaccurate measure ofmaterial dimensions, rather than molecular relaxation (andso a reduction in modulus) of the constituent tapes.

2. Experimental procedure

2.1. Penetrative falling weight impact testing

Falling weight impact testing can provide analyticalinformation about the mechanism on impact such as spec-

Fig. 1. Schematic of falling weight impact test set-up.


imen displacement, duration of impact and energy absorp-tion, but are limited to lower velocities, <10 m s�1. The bal-listic energy absorption of all-PP composite laminates hasbeen investigated by impacting laminates at much greatervelocities (>250 m s�1) and these results will be presentedelsewhere. Penetrative falling weight impact tests were per-formed as described by ASTM 5628-96 (using ASTMgeometry ‘FE’), on a range of woven all-PP compositeplates. The falling weight impact test machine used is man-ufactured by Ceast, Italy, and has a 10 mm radius hemi-spherical striker suspended at a measured height which,upon release, falls freely to strike the specimen fixed in acircular aperture beneath. A cut away schematic of theimpact aperture is shown in Fig. 1.

The impact striker contains a force transducer whichmeasures the force upon impact, and this force/time pro-vides a measurement of the energy absorbed by the impact[42]. The force/time curve is first integrated, to give the areaunder the curve, Atotal [43]:

Atotal ¼Z 1

0

F ðtÞdt. ð1Þ

And then energy absorbed by impact, Eimpact can be calcu-lated by

Etotal ¼ v0 � Atotal 1� v0Atotal

4ES

� �; ð2Þ

where Etotal = total energy of impact, v0 = impact velocity,Atotal = area under force/time curve and ES = energy ofstriker. It is possible to alter the energy of the striker byvarying the initial height or the mass on the striker. In thesetests, height is kept constant at 1 m and mass is varied inorder to vary impact energy while maintaining constant im-

pact velocity. For penetrative impact tests, the initial en-ergy is always at least twice the energy absorbed inpenetration in order to increase measurement accuracy.Energy losses due to friction in falling, sound, and heatare negligible, so will be ignored.

2.2. Non-penetrative falling weight impact testing

Non-penetrating impact tests were performed to investi-gate the mechanisms of impact damage on all-PP plates.Tests were performed on a similar Ceast falling weightimpact machine with an identical striker, but also equippedwith a mechanism to catch the striker on rebound from ini-tial impact, and prevent further strikes which occur follow-ing rebound from non-penetrating impact. The impactenergy was kept constant at 20 J. Prior to testing, speci-mens were airbrushed with a fine, random, high contrastspeckle pattern. High resolution images of these surfaceswere captured using an ARAMIS 3D strain mapping sys-tem manufactured by GOM GmbH (Germany), equippedwith two Vosskuhler CCD 1300F high resolution digitalcameras [44]. Following impact, images of the specimenswere captured again, and the ARAMIS software can pro-duce a three dimensional map of the surface deformationdue to impact damage. This requires the speckle patternto remain intact during impact, and so complete maps can-not be produced if the surface layer is damaged or the pat-tern is removed by abrasion with the impact striker orgripping mechanism. The application of digital surfacestrain mapping has been successfully applied to bi-direc-tional thermoplastic composites undergoing similar defor-mation due to stamping rather than non-penetrativeimpact [45] and has been applied to the thermoforming

0.8 1.2 1.6 2.0 2.4 2.8 3.2 3.6 4.00

20

40

60

80

100

120

140

(24)

(24)

(18)(18)

(12)(12)

(9)

(9)

(6)

Tota

l Pen

etr

ativ

e E

ner

gy [J

]

Thickness [mm]

Fabric A Fabric B PP/Woven Glass 40% wt

(...) Number of plies

Fig. 2. Absorbed penetrative impact energy vs. specimen thickness fordifferent all-PP and a commercial woven glass reinforced PP compositeshowing a non-linear relationship between specimen thickness andabsorbed energy.

0 1 2 3 4 5 60

2

4

6

8

10

12

For

ce [k

N]

Time [ms]

9 Plies (1.2mm) 12 Plies (1.5mm) 18 Plies (2.3mm) 24 Plies (3.0mm)

Fig. 3. Force vs. time curves of typical all-PP composites with varyingthickness showing an increase of impact time with 1.2 mm thick specimensindicating greater deflection during impact.


of all-PP composites elsewhere [12]. A more traditionalmethod for assessing damage in non-penetrating impactof composites is ultrasonic C-scan, but as this method relieson varying densities within a fluid medium, typically water,it cannot easily be applied for a material such as polypro-pylene with a density so close to the density of water,1 g cm�3. The acoustic damping and the lack of large scaledelaminations of tough materials such as polypropylenealso reduce the efficiency of ultrasound scanning methods.

2.3. T-peel interfacial characterisation

To determine the effect of temperature of intertapebonding of individual tapes, T-peel tests are performed inaccordance with ASTM 1876; peeling of a pair of com-pacted tapes is performed in a Hounsfield HK25S tensiletesting machine fitted with a 5 N load cell, appropriategrips and data acquisition software. The crosshead dis-placement causes two bonded tapes to peel apart at 180�in a mode I failure. The tests were performed at crossheaddisplacement of 5 mm min�1, and each test was repeated atleast five times to ensure reproducibility. The values pre-sented for peel force are defined as the force per unit widthof tape required to peel the tapes apart, since this tends to aconstant value during peeling.

3. Results and discussion

3.1. Penetrative impact performance

Due to the dimensional effect of processing conditionson the compaction of all-PP specimens, as discussed else-where [11], it is important to determine the geometricaleffect due to different specimen dimensions on the impactperformance. Fig. 7 shows the effect of specimen thicknesson penetrative impact resistance for two different all-PPcomposites composed of fabrics A and B, with a wovenglass–PP for comparison. As described in [11], fabric B iswoven from a tape with a lower draw ratio, and so greaterply thickness. This leads to a greater thickness for the samenumber of plies compared to fabric A. Fig. 2 also shows acomparison of thickness with number of plies (shown inbrackets). For specimens less than 2 mm thick, there isclearly a non-linear relationship between specimen thick-ness and penetrative energy for all-PP composites.

This is explained by examining the force–time curves ofthe impact penetrations for different thickness of all-PPcomposites. Fig. 3 shows four typical force–time curvesfor fabric A. For thicker specimens, behaviour is very sim-ilar, with maximum force increasing with specimen thick-ness. For specimens with a thickness of 1.2 mm, a greatertime to reach maximum force is observed. This is due toincreased deflection of the specimen upon impact; very thinspecimens exhibit greater flexural deformation. This willlead to a greater impact duration and a greater energyabsorption, and so very thin specimens show a dispropor-tional penetrative energy compared to their relative thick-

ness (see Fig. 4). Considering this non-linearity, allfurther impact tests are performed on specimens of greaterthan, but close to 2 mm thickness, so that thickness can benormalized in the approximately linear region of theimpact energy/thickness relationship. In any case, theimpact performance of the all-PP composites described inthis study can be considered a minimum performance, sincethis rather high thickness of all-PP composites gives thelowest impact performance.

Previous studies [12,37,39] have shown that the mechan-ical properties of all-PP composites are determined by pro-cessing parameters and this is also true of impact energyabsorption. Fig. 5 shows the effect of compaction tempera-ture and pressure on the specific tensile strength and spe-cific tensile modulus of all-PP composites. The specificproperties are defined as the property in question (i.e., ten-sile modulus or strength) divided by density. Since process-ing conditions alter composite density by affecting interplyvoid closure, it is important to isolate the effect of process-

0.8 1.2 1.6 2.0 2.4 2.8 3.2 3.6 4.00

10

20

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40

(24)(24)

(18)(18)

(12)

(12)

(9)

(9)

(6)

Tot

al P

enet

rativ

e E

nerg

y/T

hick

ness

[J.m

m-1]

Thickness [mm]

Fabric A Fabric B PP/Woven Glass 40% wt

(...) Number of plies

Fig. 4. Impact energy normalised for specimen thickness vs. specimenthickness indicating a large impact energy/thickness absorbed by thethinnest specimens.

130 140 150 160 170 1800

10

20

30

40

50

Imp

act E

nerg

y/T

hick

ness

[J.m

m-1]

Compaction Temperature [˚C]

Fabric A Fabric B PP/Woven Glass 40% wt GMT 23% wt NMT 40%

Fig. 6. Impact energy vs. compaction temperature for two all-PP fabricsand some common reinforced PP alternatives. While both all-PP fabricsperform very well in impact, fabric B shows impact behaviour which canbe controlled by compaction temperature and fabric A shows a muchmore constant impact behaviour over a wide range of compactiontemperatures.


ing conditions on mechanical properties of the constituenttapes. Fig. 6 shows the effect of compaction temperature onthe impact resistance of all-PP plates (compaction pressureis constant at 1 MPa for each all-PP specimen shown in thisgraph), compared to PP strengthened with ‘foreign’ rein-forcements: woven glass reinforced polypropylene (Twin-tex�, Saint Gobain-Vetrotex, 40% wt.), glass-matreinforced polypropylene (GMT, Symalite�, QuadrantComposites, 23% wt.) and flax fibre reinforced polypropyl-ene (NMT, 40% wt.). These traditional composites are rep-resented as horizontal lines even though impregnation/consolidation are not necessarily viable across this temper-ature window. For both all-PP fabrics, there is a generaldecrease in absorbed impact energy with increasing com-paction temperature.

0.1 1 104

5

6

7

8

9

10

11

Spe

cific

Ten

sile

Mod

ulus

[GP

a.g-1

cm3 ]

Compaction Pressure [MPa]

W160 E11

/ρ W140 E

11 /ρ

50

100

150

200

250

300

W160 σ11

/ρ W140 σ

11 /ρ

Spe

cific

Ten

sile

Str

engt

h [M

Pa.

g-1cm

3]

Fig. 5. Specific tensile properties of woven fabric A all-PP laminates vs.compaction pressure during laminate production, in which W140 andW160 refer to woven tape all-PP specimens compacted at 140 �C and160 �C, respectively, and E11 and r11 are tensile modulus and tensilestrength, respectively. This shows that specific mechanical properties arecontrolled by applied compaction pressure during laminate production.There is a negligible effect of compaction temperature in the range oftemperatures presented here. Reproduced from [11].

As the tensile strength of woven all-PP composite platesprocessed at the same compaction pressure is largely inde-pendent of compaction temperature within the range oftemperatures shown here [37], the change in absorbedimpact energy with compaction temperature can be tracedto changes in the interfacial strength. Fig. 7 combines theabsorbed impact energy of fabrics A and B with adhesivestrength determined from T-peel tests of single tapes of Aand B. These properties are clearly linked, with impactstrength being inversely related to the interfacial strengthof the tapes.

At lower compaction temperatures, the interface isweaker and so tape debonding is the preferential mecha-nism of impact failure. This delamination, together with

120 130 140 150 160 170 1800

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200

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Pee

l For

ce [

N.m

-1]


Tape A Adhesion Tape B Adhesion

10

15

20

25

30

35

40

45 Fabric A Impact Fabric B Impact

Impa

ct E

nerg

y/T

hick

ness

[J.m

m-1]

Fig. 7. Peel force vs. impact energy for two different all-PP compositesshowing the relationship between interfacial strength (determined by T-peel testing of single tapes) and impact energy absorption. It can be clearlyseen that as interfacial strength increases, impact energy absorptiondecreases.


fibrillation and tape pull-out are highly energy consumingprocesses. As compaction temperature is increased, theinterface becomes stronger and so the dominant mecha-nism of failure appears to by tape failure very locally tothe impact site. Thus at lower compaction temperatures,as the interface is weaker, the energy absorbed is lesslocalised and the absorbed impact energy is greater[9,46]; the same effect has been reported for hot com-pacted gel-spun UHMW-PE fibres composites [47]. Asin interfacial studies described elsewhere [11] anddescribed in future publications, because the effect of pro-cessing temperature of fabric A is much less than that offabric B (see Figs. 6 and 7), the effect of temperature onabsorbed impact energy is much less dramatic in fabricA than fabric B. This also makes fabric A a more easilyprocessable and consistent material for consistent impactresistance, and so is preferred for ease of processing. Ithas now been shown that impact performance can be var-ied by altering tape processing parameters such as drawratio, as well by altering composite processing. For clarity

130 140 150 160 1700

10

20

30

40

50

Ene

rgy/

Thi

ckne

ss [J

.mm

-1]


Fabric A (0.1MPa) Fabric A (1.8MPa) Fabric A (11.4MPa) PP/Woven Glass 40% wt GMT 23% wt NMT 40% wt

Fig. 8. The influence of compaction pressure and temperature on impactenergy revealing that all-PP laminates consolidated with the lowestcompaction pressure absorb the largest impact energy.

Table 3Summary of penetrative impact performance by falling mass test comparingproperties of all-PP composites depend on processing conditions)

Material All-PP processingtemperature (�C)

Conditions pressure(MPa)

Fibrefracti

All-PP 130 0.1 –150 0.1 –130 1.8 –150 1.8 –170 1.8 –130 11.4 –150 11.4 –170 11.4 –

Woven glass/PP – – 40GMT – – 23NMT – – 40

from this point forward, the results presented shall bebased on tests performed on fabric A.

As with the tensile properties, the pressure at which theall-PP composites are compacted also affects impact prop-erties. Fig. 8 shows the effect of compaction pressure on theimpact performance of fabric A, again with woven glassPP, GMT and NMT for comparison. The all-PP specimenshave been compacted at three different pressures of0.1 MPa (via a vacuum bagging processing route), 1.8and 11.4 MPa (via a hot pressing route) as described ear-lier. Here it is clear that all-PP composites can easily com-pete on impact performance with the woven glass PPtested. As compaction pressure of the all-PP compositesincreases, absorbed impact energy is reduced even whennormalised for the decreasing thickness which accompaniesprocessing at elevated pressures. This is due to an increasein interfacial strength with increasing compaction pressure,as seen before for increasing compaction temperature. Atall compaction pressures, an almost constant absorbedimpact energy is seen over the range of temperatures tested.

Due to the relatively low density of bulk PP and evenlower density of all-PP composites (see Table 3), these com-posite plates perform very well when specific impact energy(normalised for density) is taken into account. Fig. 9 showsthe specific performances of these all-PP plates comparedto woven glass PP, GMT and NMT, all normalised fordensity. From these results, all of the all-PP plates out-per-form each of the alternatives considered here (see Table 3).

The transition in failure mode can be most clearly seenin post impact specimens which have been compacted atdifferent temperatures and pressures (see Fig. 10). By com-paring an all-PP plate which has been compacted at a lowtemperature and low pressure (140 �C, 0.1 MPa), with anall-PP plate which has been compacted at a high tempera-ture and high pressure (160 �C, 11.4 MPa), the energyabsorption mechanisms of plates which exhibit the highestand lowest impact energy absorption, respectively, can beinvestigated. Fig. 10a shows an impacted specimen whichhas been compacted at low temperature and pressure(140 �C, 0.1 MPa) and shows large amounts of fibrillation

all-PP composites to alternative fibre reinforced PP composites (impact

weighton (%)

Density(g cm�3)

Penetrative energy(J mm�1)

Specific penetrativeenergy (kJ g�1 mm2)

0.61 44.8 72.90.61 44.8 73.70.72 33.0 46.10.77 28.2 36.80.89 27.7 31.20.85 21.4 25.30.86 24.0 28.00.91 25.2 27.6

1.24 28.4 22.91.07 9.8 9.21.15 4.5 3.9

130 140 150 160 1700

20

40

60

80

Spe

cific

Ene

rgy/

Thi

ckne

ss [k

J.g-1

.mm

2 ]


Fabric A (0.1MPa) Fabric A (1.8MPa) Fabric A (11.4MPa) PP/Woven Glass 40% wt GMT 23% wt NMT 40% wt

Fig. 9. Specific impact energy for different compaction pressures vs.compaction temperature (normalised for density). Since all-PP compositesconsolidated with the lowest compaction pressure also possess the lowestdensity, the specific energy absorption is far superior to the alternativereinforced PP systems compared here.


and delamination, and a higher overall energy absorption.Fig. 10b shows a similar specimen which has been pro-cessed at high temperature and pressure (160 �C,11.4 MPa) and shows a characteristic ‘star’-shaped pene-tration which is formed by tape breakage and tearing alongtape boundaries. This is a lower, more localised energyabsorption, since the deformation is limited to the immedi-ate impact site, unlike the specimen compacted at 140 �C,0.1 MPa which shows larger deformations in the plane ofthe specimen, in the area surrounding the impact site.

Ignoring frictional losses, the energy absorbed upon fall-ing weight impact, Etotal, can be considered as follows:

Etotal ¼ ET þ EI; ð3Þ

Fig. 10. Typical impact penetration damage at different compaction temperapressure (140 �C, 0.1 MPa), shows large amounts of fibrillation and delocalishigher temperature and pressure (160 �C, 11.4 MPa) shows very localised damahole.

where ET = energy absorbed by plastic tape deformation,EI = energy absorbed by interfacial failure. The energy ab-sorbed by tape deformation, ET, will be dictated by the ten-sile behaviour of the composite. Since penetration requirestape failure, the energy absorption during impact can becompared to energy absorption during tensile testing of vir-gin tape and consolidated all-PP composites. If the impactload is considered to be solely absorbed by tapes that passthrough the impact site, an effective volume of tape whichis loaded upon impact can be calculated [48].

The effective tape volume can be modelled by consider-ing the all-PP plate as stacked, unconsolidated tapes byignoring interfacial properties and crimping, and solelyconsidering tensile properties of the tape. Alternatively,to account for the effect of tape interactions, crimping,and the effects of composite processing, the tensile proper-ties of woven specimens tested in the 0�/90� direction canbe considered [12,37]. In either case, the volume loadedin tension can be considered as two rectangular sectionsrunning normal to one another with the impact site atthe centre (see Fig. 11). Although the stain rate appliedin impact (10�2–10�3 s�1) is greater than that applied inthe tensile tests (2.5 · 10�4 s�1), the tensile strength andmodulus are relatively consistent (±15%) over the strainrates seen here [11,12], and so absorbed energy is assumedto be similar.

Since impact energies are normalised for specimen thick-ness, specimen thickness can be ignored, and rather thanconsidering effective tape volume, the effective tape area,AET, can be given by

AET ¼ 2abV fV 0f ; ð4Þwhere a = striker diameter, 20 mm, b = effective tape length,60 mm, Vf = volume fraction of tape in composite �1,and V 0f ¼ volume fraction of effective material in loadingdirection. To model impact based on the tensile behaviour

tures and pressures. Specimen (a) consolidated at lower temperature anded deformation giving a circular hole, while specimen (b) consolidated atge and breakage along tape boundaries giving a characteristic star-shaped

Fig. 11. Schematic of effective area of tapes loaded in tension duringimpact.


of stacked, unconsolidated tapes, V 0f ¼ 0:5 since half of thetapes will be running normal to the direction of loading,and so make no contribution to tensile energy absorption.To model impact based on the tensile behaviour of com-pacted woven tape plies, V 0f ¼ 1, since the tensile behaviourof the woven tape composite already accounts for ineffec-tive tapes in the fabric oriented at 90� to the loading direc-tion. Table 4 shows the energy absorbed by tensile failureof the tape together with woven tape composite tensilespecimens obtained by integrating the area under stressvs. strains curves obtained during tensile testing. Combin-

Fig. 12. Illustration of the out of plane deformation of all-PP composites. Specishows large amounts out of plane deformation and tape pull through, while11.4 MPa) shows very localised damage and limited out of plane deformation

Table 4Energy absorption mechanism of all-PP composites

Specimen

Compaction temperature (�C)Compaction pressure (MPa)Energy absorbed by impact failure (per unit thickness) (J m�1)Energy absorbed by tensile failure (J m�3)Effective impact area loaded in tension (m2)Energy absorbed by tape in effective impact area (per unit thickness) (J m�1)Percentage of impact energy absorbed by tape failure

ing this data with the energy absorbed by penetrative im-pact (see Fig. 12), shows that the energy absorbed solelyby tensile failure of tapes is very close to the energy ab-sorbed by impact failure of a woven tape composite speci-men compacted 160 �C and 11.4 MPa, of which 96% can beattributed to tape failure. However, the composite platespecimen compacted at 140 �C and 0.1 MPa shows muchgreater energy absorption than can be solely attributed totensile failure of the tapes, which in this case is only 27%.Since composite plates processed under this combinationof low pressure and temperature possess much lower inter-facial strengths, it is clear that here a large amount of en-ergy is absorbed by tape pull-out or debonding. It islikely that the combination of a weak interface with stiff,strong tapes has a synergistic effect on the energy absorp-tion of all-PP composites during impact, by allowing distri-bution of impact loading to a greater volume of tape. Thefibre or tape pull-out is clearly visible from the difference inout of plane deformation of the two samples, as shown inFig. 12.

It is worth noting, that although one can optimiseabsorbed impact energy through the control of the inter-face, this may result in a composite that has such poorinterfacial strength that it may be unsuitable for somestructural applications. However, this shows that compos-ites properties can be tailored through the interface in thefinal processing step to suit the composite application. Thisis a unique feature of all-PP composites compared to tradi-tional composites which often have fibre–matrix interac-tion which cannot be altered during production.

men (a) consolidated at lower temperature and pressure (140 �C, 0.1 MPa),specimen (b) consolidated at higher temperature and pressure (160 �C,

.

Tape Compacted plate

– 140 160– 0.1 11.4– 4.5 · 104 2.5 · 104

2.3 · 107 5.2 · 106 1.0 · 107

1.2 · 10�3 2.4 · 10�3 2.4 · 10�3

2.7 · 104 1.2 · 104 2.4 · 104

– 27% 96%

Fig. 13c. Tensile surface of all-PP plate compacted at 160 �C and

Table 5Summary of non-penetrative impact specimen figures

Compactiontemperature (�C)

Impactenergy (J)

Figure

Tensile Compressive


3.2. Analysis of non-penetrative falling weight impact

damage

In order to assess the damage tolerance of all-PP com-posites, non-penetrating impact tests were performed ona range of specimens manufactured from fabric A withthree different compaction temperatures, but constant pres-sure (4 MPa). These specimens were impacted with 20 Jand the compressive and tensile plastic surfaces strainswere analysed with an ARAMIS digital optical strain map-ping system. The major surface strains on the tensile (oppo-site to impacted) side due to the impact were compared andare presented in Figs. 13a, 13b and 13c and for the com-pressive (impacted) side in Figs. 17a, 17b and 17c. The pro-cessing temperatures of specimens in these images areshown in Table 5. Some of these strain maps show discon-tinuities of strain (shown as holes in the diagrams) whichare due to either surface damage (tensile side) or removalof surface pattern due to abrasion of the striker uponimpact (compressive side). Clearly visible on all images isalso the circular pattern of the specimen clamp; this isshown as either a ring of high strain due to surface defor-

Fig. 13a. Tensile surface of all-PP plate compacted at 125 �C andimpacted with 20 J.

Fig. 13b. Tensile surface of all-PP plate compacted at 140 �C andimpacted with 20 J.

face face

125 20 13a 17a140 20 13b 17b160 20 13c 17c

mation, or by discontinuous strain mapping due toremoval of the surface pattern by the clamp. In all thestrain mapping sections detailed below, specimens are ori-ented with tapes running approximately horizontally andvertically in the plane of the page.

Figs. 13a, 13b and 13c show a range of damage processesfor the specimens compacted at 125, 140 and 160 �C, respec-tively, and subjected to a 20 J impact. As compaction tem-perature increases from 125 to 160 �C there is a transitionfrom anisotropic behaviour to a more isotropic state, andthe areas of greatest strain are located in regions at ±45�to the tape directions. This is because the tensile modulusof composites loaded in the ±45� direction is approximately40% to that of the 0� (tape) direction [37], so stress in thisdirection is much more easily relieved by deformationthrough interply shearing. This is illustrated in Fig. 14.Fig. 15 shows the stress/strain curves of two consolidatedwoven fabric plates compacted at 140 and 160 �C, and testedat 0�/90� and ±45� to the tape direction. This illustrates thedifference in moduli, but also the difference in elastic limitsof 0�/90� and ±45� specimens. It is likely that there is somestrain in the 0� and 90� direction tapes in the impacted plate,but this is elastic strain which is dissipated in the rebound ofthe striker, and not visible in the post impact specimens,whereas the strain in the ±45� direction to the tapes is plasticand so is seen in the post impact strain map.

This directional change in modulus is also affected bycompaction temperature; specimens compacted at 125 �Cpossess poor interfacial bonding and thus surface tapes willbe much freer to shear (the ‘trellis’ effect [49]), than those

Fig. 14. Stress dispersion in different angles to tape direction showing the ease of plastic deformation during loading at ±45� to tape direction.

0 10 20 30 40 500

50

100

150

200

Str

ess

[MP

a]

Strain [%]

Compaction Temperature (Loading Angle) 140˚C (+/-45˚) 160˚C (+/-45˚) 140˚C (0˚/90˚) 160˚C (0˚/90˚)

Fig. 15. Typical stress vs. strain curves for woven tape compositespecimens loaded in tension at 0�/90� and ±45� to tape direction, showinggreater plastic deformation of specimens loaded at ±45� to tape direction.

Fig. 17a. Compressive surface of all-PP plate compacted at 125 �C andimpacted with 20 J.

Fig. 16. Illustrating orientation of strain sections shown in Figs. 18 and19.


compacted at 160 �C which will be much more firmlylocked in position. Thus increased compaction tempera-tures lead to a more homogenous system.

Fig. 18 illustrates the effect of the consolidation on thestrain of the tensile surface of the laminates. The graph rep-resents the major strain as a function of the radial distancefrom the centre of the impact site (20 mm is the radius ofthe clamped region) for two sections at 0� and ±45� tothe tape direction (see Fig. 16). The specimen compactedat 125 �C shows that strain in the section radiating in the0� direction decreases with increasing distance for impactsite, as would be expected, but more importantly, muchgreater strains are seen in the section radiating at ±45� tothe tape direction. This shows the heterogeneity of straindistribution in this specimen. Contrary to this, the speci-men compacted at 160 �C shows very similar strain distri-bution whether a section is taken at 0� or ±45� to thetape direction.

Figs. 17a, 17b and 17c show specimens which have beencompacted at 125, 140 and 160 �C, respectively, and sub-jected to a 20 J impact. These show a similar situation asfor the tensile side of the impact (see Figs. 13a, 13b and13c), but now even at the compaction temperature of160 �C, the strain distribution is inhomogeneous, and iscross-shaped due to the anisotropy of the system. As com-paction temperature, and so interfacial strength, increases,

Fig. 17b. Compressive surface of all-PP plate compacted at 140 �C andimpacted with 20 J.

Fig. 17c. Compressive surface of all-PP plate compacted at 160 �C andimpacted with 20 J.

0 4 8 12 16 200

1

2

3

4

5 125˚C 20J +/-45˚ 125˚C 20J 0˚/90˚ 160˚C 20J +/-45˚ 160˚C 20J 0˚/90˚

Maj

or S

trai

n[%

]

Distance from Centre [mm]

Fig. 18. Surface strain section of typical tensile face of impact specimenshowing the change in major strain with increasing distance from impactcentre in the ±45� and 0�/90� directions for specimens processed with twodifferent temperatures.

0 4 8 12 16 20-6

-4

-2

0

2

4

6

Maj

or S

trai

n[%

]

Distance from Centre [mm]

125˚C 20J +/-45˚ 125˚C 20J 0˚/90˚ 160˚C 20J +/-45˚ 160˚C 20J 0˚/90˚

Fig. 19. Surface strain section of typical compressive face of impactspecimen showing the change in major strain with increasing distance fromimpact centre in the ±45� and 0�/90� directions for specimens processedwith two different temperatures. Clearly visible are oscillations whichrelate to the woven structure of the tape in the fabric.


the cross shape becomes less defined due to better loadtransfer between neighbouring tapes.

Like Fig. 18, Fig. 19 shows the strain distribution of 2sections on the compression sides of two specimens, com-pacted with two different temperatures, and then impactedwith 20 J. Again, a similar effect is seen, with 0� and ±45�sections having very much different magnitudes of strainfor the specimen compacted at 125 �C, but are quite similarfor the specimen compacted at 160 �C. The compressiveregion in the middle of the specimen is also seen in thisgraph as an area of negative (compressive) strain. Fromthis, it can also be seen that the plastic deformation dueto the striker at the impact site is much greater for the spec-imen compacted at 125 �C than at 160 �C. This implies thatthe specimen compacted at 125 �C, would have lower resis-tance to denting than the specimen compacted at 160 �C,despite a greater penetration resistance. This is an impor-tant consideration for applications in which the materialhas to be resistant to surface damage as well as providingpenetrative protection. The oscillations seen in the strainmagnitude for the specimen shown in Figs. 18 and 19, havea wavelength of approximately 4 mm, and this can be

explained by the fact that the strain is calculated on the sur-face of a woven fabric, with each tape being approximately2 mm wide. The peaks in the strain represent the tapes inthe 0� direction which are running radially from the impactsite in the 0� direction. The tapes which are perpendicularto these are not directly loaded and so show much lowerstrain. The specimen compacted at 160 �C shows much lessoscillation due to the increased bonding and superior loadtransfer as described above.

The use of optical strain mapping only provides infor-mation about surface strains, and since compressive andtensile faces show different strain situations, it is likely thatwhile the surface strains indicate the through-thickness

Table 6Summary of the effect of processing conditions on impact properties of all-PP composites

Property Increasing compactiontemperature or pressure

Density m

Penetrative impact resistance .

Peel strength m

Uniformity of surface strain m


strain distribution, they cannot totally describe the straindistribution in a specimen. Since the outer surfaces of thespecimen are gripped during impact, it is likely that greaterstrain is present in sub-surface plies due to interply shear-ing following delamination. Thus information obtainedby optical strain mapping must be considered as an indic-ative rather than an exhaustive description of the strain dis-tribution, and thus energy absorption of a specimen.

4. Conclusions

The impact performance of all-PP composite materialshas been analysed through penetrating and non-penetratingimpact by falling weight testing. From these results, someconclusions may be made; all-PP composite plates possessexcellent resistance to falling weight impact penetrationand can compete or outperform glass or natural fibre rein-forced polypropylene (see Table 3). The dominant failuremodes of all-PP composites in impact are delaminationand tape fracture. Since the interfacial strength of all-PPcomposites is controlled by processing conditions, it is pos-sible to tailor the impact resistance by altering tape produc-tion parameters or composite compaction parameters.Impact performance increases with decreasing interfacialstrength, with penetration energy increasing with decreas-ing compaction temperature and pressure. All-PP compos-ites which are optimised for impact, however, may notpossess adequate interfacial strengths to make them viablestructural components. Alternatively, such all-PP compos-ites may find applications as low-cost alternatives to currentrigid impact protection materials which are aimed at protec-tion from low velocity impacts. A summary of the impactproperties of all-PP composites plates is shown in Table 6.

Acknowledgements

The co-extruded PP tapes used in this study were preparedusing facilities at Lankhorst Indutech BV, Netherlands. Thiswork was sponsored by the Dutch Government’s Economy,Ecology and Technology (EET) programme for sustainabledevelopment, under Grant No. EETK97104.

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