Penetration Mechanisms in Glass Laminate-resin Structures

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Materials and Design 34 (2012) 541–551

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Penetration mechanisms in glass laminate/resin structures G.J. Appleby-Thomas a,⇑, P.J. Hazell a, R. Cleave b a b

Cranfield Defence and Security, Cranfield University, Shrivenham, Swindon SN6 8LA, UK Hamilton Erskine Ltd., 17 Moss Road, Ballygowan, Co. Down BT23 6JQ, UK

a r t i c l e

i n f o

Article history: Received 27 January 2011 Accepted 5 May 2011 Available online 13 May 2011 Keywords: A. Elastomers and rubbers B. Laminates E. Impact and ballistic

a b s t r a c t The ballistic response of composite structures comprising differing laminated float glass/polycarbonate replacement resin (PRR) elements was studied. In order to provide materials data for future modelling work, sphere-impact tests were employed to determine the high strain-rate response of the elastomeric resin. Larger-scale armour simulants comprising glass-laminate-fronted cylinders of PRR were also investigated using lead antimony-cored 7.62 mm  51 mm NATO Ball rounds in order to interrogate their behaviour under impact. Penetration mechanisms were studied via the use of high-speed video equipment. Projectile defeat in the resin was observed to depend on the degree of projectile disruption, with a greater degree of comminution leading to enhanced behaviour. This confirmed the importance of the elastomeric properties of the resin in behaviour under ballistic impact in these structures. The interaction between the glass disrupting layer and the backing absorber was found to be key to minimising subsequent penetration. The use of asymmetric float glass laminates incorporating a thinner disrupting outer surface was found to reduce subsequent depth of penetration by as much as 52% compared to similar areal density monolithic systems. High-speed video footage implied that the thinner outer layer acted to blunt the incident projectile, while the backing thick layer of glass exhibiting a Hertzian cone-like ‘‘plugging’’ failure mechanism. In addition analysis of high-speed video showed that the penetration rate in the resin was initially constant, implying penetration analogous to hydrodynamic behaviour. Ó 2011 Elsevier Ltd. All rights reserved.

1. Introduction While monolithic armour solutions, such as Al 5083-H32 [1] and polycarbonate [2] are relatively commonplace, composite systems allow more flexibility in design, with the potential to exploit useful properties of different elements in a complimentary manner [3–8]. Composite armour systems typically comprise two elements: (1) a disrupter, ideally with a hardness > that of the likely threat, designed to fracture or otherwise break up an incident projectile, and (2) an absorber, designed to dissipate the kinetic energy of an incident projectile [3,5,8,9]. In order to fully characterise such systems (e.g. for the purpose of simulation), knowledge of both material properties of the individual elements, and the way in which they interact under loading, is required. Very many novel combinations of armour have been employed. For example, Übeyli et al. [3] and Hetherington [10] both carried out similar studies investigating the ballistic response of alumina (Al2O3)/aluminium composite armours. Übeyli et al. conducted a series of experiments to compare the response of high strength low alloy (50CrV4) steel to the laminated alumina/aluminium composites (in varying configurations) following impact with ⇑ Corresponding author. Tel.: +44 (0) 1793 785731; fax: +44 (0) 1793 785772. E-mail addresses:[email protected] (G.J. Appleby-Thomas), [email protected] (P.J. Hazell), [email protected] (R. Cleave). 0261-3069/$ - see front matter Ó 2011 Elsevier Ltd. All rights reserved. doi:10.1016/j.matdes.2011.05.006

7.62 mm  51 mm M61 armour piercing (AP) rounds. The composite systems considered were shown to provide weight savings of up to 26% whilst maintaining ballistic resistance. As well as highlighting a variety of failure mechanisms in the composite, Übeyli et al. were able to show that an alumina (facing)/aluminium (backing) thickness ratio of 1:3 appeared optimal for ballistic purposes for a given armour areal density. Interestingly, via a comparison of numerical simulations and experiment (again, involving attack with 7.62 mm AP ammunition), a similar thickness ratio – in the range 1:2.5–1:2.9 – was identified by Hetherington [10] for a comparable composite arrangement. In similar work, Jena et al. [4] investigated the ballistic response to impact with 7.62 mm AP rounds of a variety of steel (both as-received and heat-treated)/ Al-7017 aluminium and vacuum-bonded steel/DyneemaÒ (a high specific strength and high stiffness polyethylene fibre-based material) composite targets. The importance of target configuration – namely maintaining a gap between the steel and Al-7017 layers – was highlighted. Steel/DyneemaÒ systems were also shown to be extremely effective, with weight savings of c.55% achievable when the ballistic response of the composite inclined at 30° was compared to that of 380 BHN rolled homogenous armour under normal attack. While useful for applications such as vehicle armour, the solutions described above [3,4,10] are not applicable to situations where a tranparent solution is required. Transparent armour is


G.J. Appleby-Thomas et al. / Materials and Design 34 (2012) 541–551

an important niche, e.g. for windows in both armoured-vehicles and buildings. However, by their nature, material choices for transparent armours are limited. Typically a composite of several elements is employed, e.g. laminates of glass and/or transparent ceramics (such as spinel or sapphire – e.g. single alumina crystals), and polycarbonate. Polycarbonate, as well as possessing an inherent degree of ballistic resistance [2], is typically employed as a backing layer to impede crack formation and catch residual projectile/comminuted target material. Inter-laminate bonds are a further important element and normally comprise a polymer film such as polyvinyl butyrate (PvB) [8,9,11]. Unfortunately, in order to meet likely threats, high laminate thicknesses are often required. This increases armour weight and in the process can lead to degradation in optical transparency [12]. Reduction of the areal density of transparent armours either by optimising the architecture of current solutions or via the inclusion of more weightefficient materials solutions is therefore desirable. Polyurethane, an elastomeric thermosetting resin, is a current candidate material for transparent armour systems. A transparent polyurethane-based armour system, known as CleargardÒ, has been successfully developed and brought to market by BAE Systems [13,14]. This material has, for a given areal density, shown superior ballistic performance compared to both polycarbonate and acrylic and has been shown effective against threats ranging from fragmentation (in a monolithic layer) to Pb-cored 7.62 mm ammunition (as part of a composite solution). A similar material, from an alternate source, is polycarbonate replacement resin (PRR); a polyurethane resin [8]. PRR is a potential candidate for nano-reinforcement [15] and has a similar refractive index to glass, rendering it transparent. Its extended cross-linking imparts a high toughness, while the material has a lower density then glass, providing a direct weight saving if substituted for other elements of a transparent armour system. When uncured PRR is a viscous liquid – a property which allows it to be cast into complex geometries. Usefully, if cured in contact with glass a strong chemical bond is formed. While suitable for the energy-absorbing element of transparent armour, the low hardness of PRR means that a hard-faced disruptor is required on the face adjacent to the threat. Krell et al. [11] considered ceramics such as spinel and alumina for this application. Their high hardness (typically HV10 > 20 GPa) provides maximal ballistic strength. Following a discussion of manufacturing routes leading to the production of a variety of different configurations of spinel and alumina, Krell et al. carried out a series of ballistic tests using 7.62 mm  51 mm AP ammunition which highlighted the importance of using the lowest practical grain size. Not only does this act to limit the potential for lattice-based elements of microplasticity (e.g. dislocations and twins) to migrate within the structure, but smaller grain sizes also result in increased optical transparency. Klement et al. [9] studied a range of potential facing materials including float glass, glass ceramic, quartz glass, ALON (aluminium oxynitride) and sapphire. Depth of penetration (DOP) tests using both 7.62 mm  51 mm WC–Co cored AP8 (NAMMO) and 7.62 mm  54 mm R B32 API hard steel cored (RAPI) AP projectiles were carried out on single materials as well as composites. These composite systems comprised glass and ceramic-faces with float glass central sections backed by polycarbonate layers. Sapphire (HV0.1 = 2158) and ALON (HV0.1 = 1772) were shown to reduce DOP by 22% and 31% respectively compared to float glass (HV0.1 = 572) using NAMMO projectiles and by 57% and 38% respectively using RAPI projectiles (in the latter case compared to glass ceramic – HV0.1 = 633). Similar results were found when comparing the performance of the glass and sapphire faced composites considered. Overall, these results confirmed the advantages of increased hardness in defeating incident projectiles. As discussed above, many authors have considered individual elements of transparent armour systems [9,11]. However, as sug-

gested by the experimental and modelling work on glass-faced PRR-backed systems undertaken previously by Hazell et al. [8], the interaction between the various elements is also of importance. Hazell et al. showed that following impact with a lead antimonycored 7.62 mm  51 mm bullet, significant projectile disruption occurred during penetration, although interestingly the shape of the resultant penetration cavity was largely governed by the elastomeric nature of the PRR. Analysis of high-speed video of the penetration event plus recovered material – as well as comparison to computational simulations – indicated that both core and jacket material were deposited along the penetration path, with the bullet essentially turning inside-out. In essence, these results showed that at elevated strain-rates (e.g. at the tip of an impacting projectile), materials often behave in a hydrodynamic manner – i.e. strength effects become negligible [16]. The material response in such strain-rate regimes is largely governed by the equation-of-state, which allows pressure, energy, shock velocity, particle velocity (the continuum velocity of atoms or structures propagating a shock) and density to be related to each other. The authors of this paper previously experimentally derived the high strain-rate equationof-state for PRR [17] via the impedance-matching technique [16,18,19]. At high particle velocities (uP)/impact pressures good agreement with the behaviour of polyurethane was observed. A linear relationship was apparent in the particle velocity–shock velocity plane up to uP  0.4 mm/ls. However, below this value (and particularly for uP < 0.2 mm/ls), the experimental data was observed to trend below the equation-of-state of polyurethane. Such non-linearity has been attributed elsewhere to disparity in the strength of the backbone (polymer chain) and inter-chain forces, with the latter an order of magnitude smaller than the former, leading to a multi-stage response to compression [20]. Such information, when combined with a suitable strength model, can aid numerical simulation of impact phenomena. However, in the ballistic impact regime strength and failure mechanisms are of greater importance. Consequently, in this paper the failure and lead-cored projectile defeat mechanisms of a float glass/PRR-based transparent armour concept are considered. This PRR-based solution was conceived in an attempt to reduce the areal density of such armours by c.20%. This work builds on the computational and experimental studies undertaken on similar systems by Hazell et al. [8]. The overall aim was to extend understanding of the mechanisms controlling both penetration into PRR and, unlike the pure float glass considered in Ref. [8], the effects and projectile-defeat mechanisms associated with laminated float glass disrupting layers. To this end, experiments have been undertaken to increase understanding of both the properties and interaction of the various elements of this composite armour system; preliminary results are reported here. Two different approaches were adopted: (1) sphere-impact tests [21] to study the deformation mechanisms of projectiles within the resin itself, with both deforming and non-deforming projectiles employed to maximise the extent of information gained, and (2) ballistic impact tests involving single 7.62 mm  51 mm NATO Ball rounds impacting the centre of float glass-faced PRR targets to study, by comparison to the sphere impact results, penetration mechanisms into the composite system.

2. Experimental Two different types of tests were employed to investigate different aspects of the ballistic response of glass/PRR composite structures. Sphere-impact tests were used to investigate the nature of penetration into the PRR-only. Ballistic-impact tests were also conducted; these used 7.62 mm ammunition on an indoor firing range to try and clarify the effect of differing glass-laminates. An

G.J. Appleby-Thomas et al. / Materials and Design 34 (2012) 541–551

outline of the experimental approach adopted for each type of test is presented in Sections 2.1 and 2.2 respectively.


camera operating at a maximum frame rate of 120,171 frames/s and typical exposure of 8 ls. 3. Results and discussion

2.1. Sphere-impact Sphere-impact tests were conducted using a Ø30 mm smoothbore, 5 m barrel, single-stage gas-gun. Both deforming Pb and more rigid (essentially non-deforming) WC–Co projectiles were selected. These materials were selected as they represent typical standardissue bullet core materials. Impact velocities were chosen such that both projectiles possessed similar kinetic/impact energies of c.1.8 kJ. Full details of the projectiles used and the associated impact conditions are set out in Table 1. Projectiles were encased in acetal sabots which were stripped immediately on exit from the muzzle. Targets comprised polycarbonate boxes (with 6 mm thick sides) assembled to provide sufficient internal volume to cast 100 mm  100 mm (impact face), 95–100 mm deep, PRR in situ. These targets were placed c.20 mm from the exit of the sabot stripper, with impact phenomena recorded using a Phantom V12 high-speed camera. Frame rates of 15,000–22,500 frames/s were employed with typical exposure times of 10 ls. Further via. the use of fiducial markers these recordings also allowed calculation of the projectile impact velocities. In addition to information obtained from high-speed video, the final depth of penetration and other key features were assessed by visual analysis of the targets post-impact.

2.2. Ballistic impact Ballistic testing using 7.62 mm  51 mm NATO Ball rounds was undertaken using a proof barrel located 10 m from the targets front face. A single bullet was fired into the centre of each target. Two different geometries of target were employed: (1) outer Ø170 mm tubes, and (2) outer Ø120 mm tubes, in both cases comprising 4 mm thickness polycarbonate, filled with PRR and faced with varying architectures of float glass laminates. Target assembly was conducted by Hamilton Erskine Ltd. (N. Ireland, UK) and took advantage of the ability of PRR to chemically bond to glass on setting. Essentially, the required polycarbonate tube was placed onto the rear surface of the face-down prepared laminate before being filled with resin. Further, to allow for resin cure times, the PRR was typically cast in 3–4 layers, each of 30–50 mm thickness, with successive layers only cast once the previous one had set. Due to the proprietary nature of the laminate combinations employed specific details of the float glass architectures tested are withheld. Instead, only the generic nature of the facing glass arrangements are discussed. Further, specific results are only compared for similar areal density systems – with a maximum deviation of +25% from the monolithic layer employed as a standard. None-the-less, it is felt that this level of detail is sufficient to highlight key mechanisms controlling transparent armour defeat in such systems. PvB interlayers were employed between laminate layers unless otherwise stated. Bullet velocities were measured using a combination of a Doppler-based measurement system and a series of light-gates known as a ‘sky screen’. An average impact velocity of c.812 m/s was recorded, in good agreement with BR6 standards [22,23]. Impact conditions were monitored using a Phantom V12 high-speed

3.1. Sphere-impact As expected, the morphology of penetration into PRR targets was found to depend upon the choice of projectile. With nondeforming WC–Co projectiles, a post-impact penetration path was formed within the resin. This subsequently closed behind the intact projectile. Interestingly, analysis of high-speed video footage showed that the final depth of the projectile within the target was significantly less than the maximum depth of penetration. Both the closure of the penetration path and subsequent bulk recovery appear to be linked to the elastomeric nature of the resin. Both of these phenomena are clearly visible in Fig. 1. Fig. 1 shows the penetration of a WC–Co projectile into a typical resin target. Such results suggest that deformation of the resin leads to the build-up of stored elastic strain energy. Once the sphere had begun to penetrate the resin the penetration path began to close behind. Finally, once a maximum depth of penetration had been reached, relaxation of the elastically deformed resin resulted in a notably lower final depth (e.g. the frame at 1558 ls in Fig. 1 as opposed to that at 356 ls). The significantly softer Pb projectiles, however, deformed on impact/during subsequent penetration. Analysis of both the targets and high-speed video captured during the impact indicated that projectile material was continually broken up during passage through the resin. This disrupted material was pushed to one side as the projectile passed and was subsequently trapped in situ when the deformed resin relaxed, leaving behind a series of characteristic ‘wings’ which pointed along the direction of formation of the penetration path (observed previously in PRR by Hazell et al. [8]). These structures are discernable in Fig. 2 which shows the penetration of a Ø12.5 mm Pb projectile into a pure PRR target. Again, relaxation similar to that seen with the WC–Co projectile shown in Fig. 1 was apparent after the projectile had reached its maximum depth of penetration. This suggested that a large part of the penetration process is controlled by the elastomeric properties of the resin. Further, the lower overall depth-of-penetration compared to the WC–Co case shown in Fig. 1, despite the similarity

Fig. 1. Selected high-speed video frames showing the penetration of a Ø12 mm WC–Co sphere into a typical resin target at 515 ± 10 m/s.

Table 1 Key properties of projectiles employed for sphere impact tests. Projectile

Mass (g)

Vimpact (m/s)

Impact energy (kJ)


Ø12.0 mm Tungsten carbide (WC–6Co) Ø12.6 mm Lead (Pb)

13.5 11.7

515 ± 10 550 ± 10

1.79 1.77

Atlas Ball & Bearing Co. Ltd., Walsall, UK (grade 25) G.E. Fulton & Son, Bisley Camp, Surrey, UK


G.J. Appleby-Thomas et al. / Materials and Design 34 (2012) 541–551

Fig. 2. Selected high-speed video frames showing the penetration of a Ø12.5 mm Pb sphere into a PRR target at 550 ± 10 m/s.

Fig. 3. Frames extracted from high-speed video showing the impact of an 822 m/s 7.62 mm NATO Ball round into a thinner-fronted bi-layer asymmetric float-glass faced PRR backed cylindrical target.

in impact energies (Table 1), suggested that the area of material presented to the resin by the penetrating projectile was important.

Essentially a larger incident projectile surface area (caused by projectile disruption) appeared to dissipate the projectiles kinetic


G.J. Appleby-Thomas et al. / Materials and Design 34 (2012) 541–551 Table 2 Ballistic impact experimental results – cylindrical laminated glass-fronted PRR-backed targets impacted in the centre by a single 7.62 mm  51 mm NATO Ball round. Exp. no.

Glass arrangement (front ? back)

Facing areal density relative to monolithic layer

Depth-of-penetration (mm) Vimpact (m/s)

Final (meas.)

Max. (HSV)

Final (HSV)

Elastic recovery

1 2 3 4 5 6 7 8 9 10 11

Thin/thick 50:50 bi-layer Monolithic 50:50 bi-layer Monolithic Thin/thick Thin/thick 50:50 bi-layer 50:50 bi-layer Thin/thick Thin/thick 50:50 (tri-layer) Thin/thick 50:50 (tri-layer) Thin/thick Thin/thick

+25% = = = = = = = = +25% +25%

822 836 813 806 838 814 811 810 813 812 818

21.1 68.2 60.5 61.1 – – 58.1 – – – –

32.4 71.2 61.8 64.2 65.9 63.0 60.9 69.6 67.8 35.2 60.8

23.9 68.2 59.0 60.5 64.0 60.8 58.8 67.4 64.8 29.7 57.4

8.5 3.0 2.7 3.7 1.9 2.3 2.2 2.2 2.9 5.5 3.4







+25% +25%

814 815

– –

36.0 44.5

32.3 42.9

3.8 1.6

12 13 14

energy over a larger surface area (and corresponding backing volume) of PRR. This allowed the projectiles energy to be dissipated over a shorter penetration distance within the PRR (albeit with a wider area of damage across the targets face), leading to a reduced subsequent depth-of-penetration.

3.2. Ballistic impact A series of frames extracted from a typical high-speed video recording of an impact into a cylindrical target are shown in Fig. 3. The transparent area to the left of each frame is the PRRfilled polycarbonate tube, with the glass facing layer visible as dark bands towards the centre. This test involved impact of a single 7.62 mm NATO Ball round, visible to the right of the target at 0 ls, into the centre of a cylindrical PRR-filled polycarbonate tube faced by a bi-layer asymmetric float-glass laminate (with a thinner outer face). Just after impact at 41 ls comminuted glass is visible flowing away from the impact site. By 83 ls the flow of ejected material has increased in volume and an area of disruption has become apparent within the backing resin. Subsequent studies at higher frame-rates showed that the width of this area grows before necking (apparent at 124 ls in Fig. 3) occurs and the area of disrupted material forms two different regions. It is postulated that at this point stored elastic strain energy imparted to the resin by the bullet impact has reached a sufficient level to allow the resin to recover about the point of greatest instability with in the flow of disrupted material. This lateral recovery/necking, analogous to the closure of the penetration path behind WC–Co and Pb projectiles observed in the sphere-impact tests (Figs. 1 and 2 respectively), continues as penetration increases. One particularly important observation is the large degree of recovery apparent along the axis of penetration. Peak penetration occurs at around 166 ls, reaching a depth behind the glass face of c.32.4 mm. However, by 373 ls recovery of at least 5 mm has occurred, with subsequent recovery eventually reducing the final depth-ofpenetration to c.24 mm. Again, this phenomenon is likely to be linked to release of stored elastic strain energy within the resin, showing that the elastomeric properties of the resin are key to the manner in which it resists penetration. Experimental results from the trial shown in Fig. 3 (test no. 1) and the first set of tests subsequently conducted are summarised in Table 2. Key parameters recorded include bullet impact velocities and depths-of-penetration, measured as required either from high-speed video footage or directly from recovered targets. As discussed, due to the proprietary nature of the glass-facing architec-

tures used only generic details are included. Here, the number of laminated layers employed, their relative symmetry and the relative magnitude of their areal density compared to the monolithic system employed in tests 3 and 5 are noted. The data presented in Table 2 strongly suggested that the thin layer-fronted asymmetric systems considered in tests 1, 6–7 and 10 exhibited an enhanced ballistic response. Consequently, the decision was taken to test a further series of these targets, although with differing inter-glass laminate material composition and thickness. While interlayer thickness appeared to have a slight effect on depth-of-penetration (at most optimisation led to a reduction of c.20% compared to a similar system with a radically different interlayer thickness) – a result encountered elsewhere [22] – discussion of the effects of this interlayer thickness is considered beyond the scope of this paper. Never-the-less as it represents a useful comparison to results from the initial tests, experimental data from these tests is also included here in Table 3. 3.2.1. Penetration rate in the PRR Following passage through the facing glass laminates, the position of the head of the region of disturbance within the PRR was monitored using captured high-speed video footage (e.g. Fig. 3) for a number of the ballistic impact tests detailed in Table 2. These tests involved a variety of different PRR-backed glass-laminate faced targets. Fig. 4 shows the resultant variation in displacement with time. The most important observation from this figure was that initial penetration largely occurred at a constant velocity (e.g. the gradient of the position–time curve was constant). This implied that the resin initially behaved in a fluid-like manner, exhibiting minimal/no strength, suggesting that a hydrodynamic treatment of the resin under impact would be valid [16,17]. Interestingly, this constant-velocity penetration occurred independent of the architecture of the target. Following the initial hydrodynamic penetration in Fig. 4, a sudden cusp is apparent as stored elastic strain energy suddenly slows the penetrating projectile/ material down. This is followed by a negative gradient indicative of contraction before a constant line of zero gradient is established indicating that the final depth-of-penetration has been reached. This penetration/contraction morphology ties in with the postpenetration recovery of the resin highlighted in Fig. 3, discussed in more detail at the beginning of Section 3.2. 3.2.2. Laminate composition As shown in Table 2, a variety of different laminate compositions were considered. These were chosen such-that glass and


G.J. Appleby-Thomas et al. / Materials and Design 34 (2012) 541–551

Table 3 Ballistic impact experimental results – cylindrical thin glass-fronted thin/thick laminated PRR-backed targets, with a variety of different glass interlayers configurations, impacted in the centre by a single 7.62 mm  51 mm NATO Ball round. Exp. no.

15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37


0.38 mm 0.38 mm 0.76 mm 0.76 mm 1.14 mm 1.14 mm 1.52 mm 1.52 mm 1.52 mm 1.52 mm 2.28 mm 0.38 mm 0.38 mm 0.76 mm 0.76 mm 1.14 mm 1.14 mm 1.52 mm 1.52 mm 1.52 mm 1.52 mm 2.28 mm 2.28 mm

PvB PvB PvB PvB PvB PvB PvB PvB SentryglassÒ SentryglassÒ SentryglassÒ PvB PvB PvB PvB PvB PvB PvB PvB SentryglassÒ SentryglassÒ SentryglassÒ SentryglassÒ

Facing areal density relative to monolithic layer

Depth-of-penetration (mm) Vimpact (m/s)

Max. (HSV)

Final (HSV)

Elastic recovery

+25% +25% +25% +25% +25% +25% +25% +25% +25% +25% +25% = = = = = = = = = = = =

804 808 812 813 800 811 816 807 796 808 806 809 805 805 805 809 810 809 807 807 809 802 804

30.5 38.7 33.7 30.5 36.8 38.7 29.8 31.7 32.8 27.9 35.6 50.3 43.8 52.5 57.1 59.3 48.9 54.9 52.9 55.6 52.7 50.8 51.4

24.1 36.2 28.6 24.1 33.1 36.2 23.5 22.2 27.8 19.7 29.8 49.3 42.6 51.4 55.9 57.1 48.3 52.7 52.1 50.5 48.3 48.3 48.9

6.4 2.5 5.1 6.4 3.7 2.5 6.3 9.5 5.0 8.2 5.8 1.0 1.2 1.1 1.2 2.2 0.6 2.2 0.8 5.1 4.4 2.5 2.5

Fig. 4. Variation of depth-of-penetration with time for a single 7.62 mm NATO Ball round penetrating into PRR behind various glass laminate faces.

laminate thicknesses were indicative of those expected for commercially-relevant bullet-resistant window architectures (e.g. Ref. [8]). From this table it is immediately apparent that targets 1, 10, 13 and 14 – namely those with a bi-layer asymmetric laminate with a higher areal density than the monolithic standard – performed between 55% and 70% more effectively than the other samples. The key differences between the response of these targets and the other arrangements considered, based on Table 2, appeared to be a substantial increase in post-impact elastic recovery. E.g. recovery was observed to be consistently up to 8.5 mm, compared to consistently (+25%) 35.2 29.6

PRR when a larger surface area of disrupted material is presented to the resin. This strong correlation between depth-of-penetration and W confirms the previous hypothesis that inter-laminate delamination (e.g. Fig. 6) as well as projectile erosion by the front glass face are not the only energy-dissipating processes operating to enhance ballistic resistance. While there is a degree of scatter (not un-expected given the nature of ballistic tests), the lowest measured depths-of-penetration, corresponding to the largest values of W, occur for the high areal density asymmetric bi-layer laminates. A significant reduction is also apparent where the asymmetric bi-layer laminates of the same areal density as the monolithic systems investigated were employed. Overall, for the non bi-layer non-thin-fronted laminate systems an average maximum depth-of-penetration of 64.9 mm was found. This compares to average maximum depths-of-penetration of 35.2 mm and 54.4 mm for the asymmetric thin-fronted bi-layer architectures with a higher and lower areal density respectively. A better point-of-comparison would be with the ballistic response of the monolithic layer float-glass targets. Consequently, using the data from Tables 2 and 3, the final and maximum depths-of-penetration (averaged across all tests with similar targets) of the two asymmetric bi-layer thin-fronted laminate architectures considered are compared to those for the standard monolithic layer employed in Table 4. Table 4 shows that replacing a monolithic glass layer with a thin front-layer asymmetric bi-layer laminate led to improvements in maximum/final depths-of-penetration of 14.9/16.8% respectively for no change in areal density. Where similar composition laminates with a thicker backing layer were employed improvements in ballistic resistance of 44.9% (maximum) and 51.9% (final) resulted compared to the monolithic arrangement, albeit accompanied by an increase in areal density. In this latter case the increase in areal density is found to more-than outweighed by the substantial increase in ballistic resistance (as discussed, absolute values are not quoted due to the proprietary nature of the arrangement considered). 4. Conclusions The ballistic response of the different elements of a proposed glass-faced PRR-based transparent armour solution have been investigated using a combination of sphere and ballistic impact tests. High-speed video footage of sphere impact experiments allowed interrogation of the influence of the elastomeric properties of PRR on penetration. In addition, high-speed video analysis of ballistic impact tests helped identify defeat mechanisms in both float glass disrupting layers and composite glass-faced PRR-backed targets. While the proprietary nature of the approach meant that exact details of the proposed glass-facing solution were not provided, the effect of modifying the glass facing by splitting it into a thin-fronted asymmetric bi-layer configuration was considered. Further, the importance of the elastomeric resin in terms of absorbing incident projectile energy was highlighted. The key conclusions which have emerged from the discussion of the experimental results presented here are highlighted below:

(1) The importance of monitoring impacts using high-speed video footage was highlighted as the final depth-of-penetration in ballistic targets is significantly less than the maximum depth due to elastic recovery in the PRR. (2) High-speed video footage suggested that PRR behaves predominantly hydrodynamically under impact, with strength-effects only becoming important once the maximum depth-of-penetration of an impacting projectile is reached (e.g. when elastic PRR recovery begins to occur). This result does not appear to the authors’ knowledge elsewhere in the literature. (3) Energy dissipation, whether via inter-laminate delamination or enhancing the surface area of contact between the penetrating projectile and the PRR, is key to minimising depth-ofpenetration – highlighting the importance of the elastomeric properties of PRR. (4) Disrupting glass-laminates comprising a thin impact face backed by a thicker glass layer were found to reduce the depth-of-penetration compared to monolithic glass plates by up to c.52%, with a small increase in areal density. This approach therefore appears to suggest that reductions in the areal density compared to current transparent armour solutions, combined with retention of the desired level of protection, are feasible. (5) Thin/thick laminate systems were shown to maximise projectile disruption and minimise subsequent depth-of-penetration via a two-stage mechanism involving (1) projectile erosion in the facing thin layer, followed by (2) plug-formation in the thicker rear glass layer. Acknowledgements The authors would like to acknowledge provision of funding by Hamilton Erskine Ltd., N. Ireland. In addition, Gareth ApplebyThomas would like to recognise the contribution of his wife, Caroline Jane Appleby-Thomas, during the final stages of paper preparation. References [1] Appleby-Thomas GJ, Hazell PJ, Millett J, Bourne NK. Deviatoric response of an armour-grade aluminium alloy. Shock Compress Condens Matter 2009:533–6. [2] Wright SC, Fleck NA, Stronge WJ. Ballistic impact of polycarbonate – an experimental investigation. Int J Impact Eng 1993;13(1):1–20. [3] Übeyli M, Orhan Yildirum R, Ögel B. On the comparison of the ballistic performance of steel and laminated composite armours. Mater Des 2007;28:1257–62. [4] Jena PK, Ramanjeneyulu K, Siva Kumar K, Balakrishna Bhat T. Ballistic studies on layered structures. Mater Des 2009;30:1922–9. [5] Hazell PJ. Ceramic armour: design and defeat mechanisms. 1st ed. Canberra, Australia: Argos Press; 2006. [6] Hazell PJ, Roberson CJ, Moutinho M. The design of mosaic armour: the influence of tile size on ballistic performance. Mater Des 2008;29:1497–503. [7] Özsßahin E, Tolun S. Influence of surface coating on ballistic performance of aluminium plates subjected to high velocity impact loads. Mater Des 2010;31:1276–83. [8] Hazell PJ, Edwards MR, Longstaff H, Erskine J. Penetration of a glass-faced transparent elastomeric resin by a lead–antimony-cored bullet. Int J Impact Eng 2009;36:147–53. [9] Klement R, Rolc S, Mikulikova R, Krestan J. Transparent armour materials. J Eur Ceram Soc 2008;28:1091–5. [10] Hetherington JG. The optimization of two component composite armours. Int J Impact Eng 1992;12(3):409–14. [11] Krell A, Klimke J, Hutzler T. Advanced spinel and sub-lm Al2O3 for transparent armour applications. J Eur Ceram Soc 2009;29:275–81. [12] Sturrock AJ. Investigation of protective glass laminates. MSc thesis. Shrivenham Campus: Cranfield University; 2003. [13] Aircraft Armour Systems (BAE Systems publication) [cited 14 04 11]. . [14] Cleargard transparent polymer (BAE Systems publication) [cited 14 04 11]. .

G.J. Appleby-Thomas et al. / Materials and Design 34 (2012) 541–551 [15] Chin SJ. Investigation of nanotechnology enhanced polyurethane replacement resin (PRR) as a transparent armour material. MSc thesis. Shrivenham Campus: Cranfield University; 2004. [16] Meyers MA. Dynamic behaviour of materials. New York: John Wiley & Sons Inc.; 1994. [17] Appleby-Thomas GJ, Hazell PJ, Stennett C, Cooper G, Cleave R. The dynamic behaviour of a modified polyurethane resin, DYMAT 2009. In: 9th International conference on the mechanical and physical behaviour of materials under dynamic loading; 2009. p. 1081–87. [18] Marsh SP. LASL shock Hugoniot data. University of California Press, Ltd.; 1980. [19] Vignjevic R, Bourne NK, Millett JCF, Vuyst TD. Effects of orientation on the strength of the aluminum alloy 7010-T6 during shock loading: experiment and simulation. J Appl Phys 2002;92(8):4342–8.


[20] Munson DE, May RP. Dynamically determined high-pressure compressibilities of three epoxy resin systems. J Appl Phys 1972;43(3):962–71. [21] Hazell PJ, Appleby-Thomas G. A study of the energy dissipation of several different CFRP-based targets completely penetrated by a high velocity projectile. Compos Struct 2009;91:103–9. [22] López-Puente J, Arias A, Zaera R, Navarro C. The effect of the thickness of the adhesive layer on the ballistic limit of ceramic/metal armours. An experimental and numerical study. Int J Impact Eng 2005;32:321–36. [23] Lawn BR. Indentation of ceramics with spheres: a century after Hertz. J Am Ceram Soc 1998;81(8):1977–94.

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