On the ballistic impact response of microbraid reinforced polymer composites

Abstract This paper presents the behaviour of microbraid reinforced polymer composites (mBRPC) subjected to impact loading conditions. Ballistic impact tests were performed by firing 7.94 mm steel balls onto composites reinforced with microbraids having different architectures, braid angles and of different materials (Kevlar® and Dyneema®). Two high speed cameras were employed to record the impact events. Experimental results revealed an improvement in the ballistic limit, of up to 19.5% for certain types of mBPRC, with respect to composites made with unidirectional fibres. Visual inspection of the impacted laminates revealed similar deformation mechanisms for composites reinforced with microbraids having different architectures and of different material. The slippage of the impactor through the layers of the laminates could have had detrimentally affected the ballistic properties of the manufactured composites. Modifications in the arrangement of the reinforcing phase are needed to fully exploit the potential of the microbraids in polymeric structures.


Introduction
High performance polymeric fibres, such as ultra high molecular weight polyethylene (Dyneema Ò , Spectra Ò ), aramids (Kevlar Ò , Twaron Ò ), LCP (Vectran Ò ) and PBO (Zylon Ò ) are widely exploited in applications in which high levels of energy absorption and protection are required. Dry fibres are mainly used in the construction of ropes, lines and nets, fabrics composed of such fibres are used in high performance textiles and vests, whilst impregnated fibre or fabric systems in hard components for ballistic applications and containment, for example. It has been demonstrated that tailored yarns can be developed by braiding fibres (for example [1][2][3][4][5][6][7]). It has also been demonstrated in [7] that the tensile properties of certain types of fibres can be enhanced by braiding; typically the tensile strength, strain to failure and energy absorption can be modified by braiding. There is sparse documentation in the open literature on the behaviour, under ballistic impact conditions, of braid reinforced polymer composites. Haijun et al. [8] investigated the ballistic properties of triaxial braided composites by firing titanium alloy cylindrical projectiles and blade-shape projectiles with the same mass of 17.5 g. Composites were manufactured using flattened 0 ± 60 triaxial carbon braid tubes and epoxy resin via an RTM process. The manufactured laminates were 10 mm thick, from which 100 mm Â 100 mm impact test coupons were cut. The ballistic limit of the braided composites were above 195 m/s and 207 m/s when impacted with flat cylinders and bladelike projectiles, respectively. Triaxial braided composites showed a higher ballistic limit and smaller damaged area with respect to satin woven composites having approximately the same areal density and fibre volume fraction, when tested under the same impact conditions. Roberts et al. [9] performed ballistic impact tests to investigate the response of 2D triaxial braided composites. An extensive series of impact tests were conducted on braided composite plates, of dimension 610 mm x 610 mm and with a nominal thickness of 3.2 mm, and braided reinforced half-rings specimens, which replicate the shape of fan cases for jet engines. Composites were manufactured using 6 layers of carbon triaxial braids having a bias angle of 0 ± 60 and epoxy resin via RTM and impacted at different speeds using soft gelatine projectiles. Fracture propagated along the bias fibre direction. However, the damage area was very localised around the impact point and no delaminations were observed. The penetration threshold was determined to be 155 ± 5 m/s and 135 ± 3 m/s for the flat plates and the half-ring specimens, respectively. The authors highlighted the importance of performing impact tests on specimens reflecting their actual shape design. This is to highlight characteristic features which otherwise could be missed using different geometries. Binienda and Cheng [10] performed impact tests on triaxial braided reinforced composites with different bias angles. The materials and manufacturing method were the same as [9]. They noted that the damage pattern and velocity penetration threshold depended on the braid architecture. Composites having 0 ± 45°bias fibres have been shown to have the best ballistic performances with respect to those having fibre orientation 0 ± 60°. However, the axial fibre content of the former specimens had twice the fibre content in axial direction with respect to the latter.
To the authors best knowledge, no information is available in the open literature on the ballistic impact response of polymer composites reinforced with high performance microbraids. It has been shown in our previous study [7] that the mechanical performance of polymer composites reinforced with microbraids can be superior with respect to those made of unidirectional fibres in terms of energy absorption, tensile strength and strain to failure. In this paper, the ballistic performance of sixteen different types of microbraid reinforced polymer composites (mBRPC) are experimentally investigated through a series of impact tests. Results are compared with those obtained from testing conventional cross-ply laminates made of unidirectional fibres (having similar areal density and fibre volume fraction, and manufactured via the same technique), deemed as the baselines.

Materials
Three high performance polymeric fibres were used in this investigation, namely Dyneema Ò SK75, Dyneema Ò SK76 and Kevlar Ò 49. Physical properties of these materials, provided by the manufacturers [11][12][13], are listed in Table 1. Fibre diameters were determined by analysis of images from scanning electron microscope (SEM).

Manufacture
The manufacture and mechanical characterisation of dry microbraids and mBRPC used in this work were described in detail in [7], and it can briefly summarised as follow.
Different types of microbraids were manufactured using the yarns having the smaller linear density by a Herzog RU2-16/80 vertical braiding machine. Microbraids having different bias angles a and architectures were created by varying the cogwheel ratio and the number of working carriers, respectively. Core-filled microbraids were produced by overbraiding a unidirectional (UD) yarn with eight bias yarns of the same material in a diamond fashion. The diameter of the microbraids and their bias angles were determined by analysis of SEM images. The microbraid linear densities were determined according to the ASTM D1577-07 [14] Standard Test Methods for Linear Density of Textile Fibers. The investigated braid patterns are sketched in Fig. 1. Physical properties of the manufactured microbraids are listed in Table 2.
The tensile properties of dry microbraids were assessed through a series of tensile tests performed on specimens with a gauge length of 250 mm and tested at a strain rate of 0.01 s À1 .
The dry microbraids were aligned in a unidirectional fashion over a spinning plate via a robotised filament winding system and subsequently consolidated into prepregs using a thermoplastic resin film (Rayofix TP) via a hot-pressing technique. Then, the prepregs were manual assembled in a cross-ply laminate and hotpressed into the final laminate form.
In order to directly compare the properties of the mBRPC with cross-ply laminates made with UD fibres and manufactured via the same route, composites having similar areal density and fibre volume fraction were manufactured using the coarser yarns. Physical properties of the mBRPC tested under ballistic impact conditions are presented in Table 3.
Given below is the nomenclature or classification used for the microbraids and mBRPC. A generic dry microbraid and mBRPC will belong to the class ''w X Y Z", where: w will be the physical form of the material, in particular ''b" stands for dry microbraids and ''c" for microbraid reinforced composites; X will be the material of the microbraid, in particular D for Dyneema Ò SK75 and K for Kevlar Ò 49; Y will denote the braid angle, where A < B < C . . .; Z will represent the braiding architecture, in particular ''1" for diamond 1/1, ''2" for regular 2/2 and ''H" for core-filled microbraids; Composites cXUD were manufactured using UD yarns; Table 1 Physical properties of the investigated materials [11][12][13]. Impact test specimens with dimensions of 150 mm Â 150 mm were cut from the manufactured laminates using a waterjet cutter.

Testing Method
The ballistic impact tests were performed in the Department of Aeronautics at Imperial College using a 50 mm calibre single stage gas gun (Fig. 2).
The gas gun has two 10 lt pressure vessels in which the launch gas can be pressurised up to 330 bar. The projectile was positioned on the front of a polypropylene sabot (Fig. 3a) and inserted into the breech of the gas gun. A breech plug was then inserted and screwed closed, positioning the sabot into the precise firing position. The specimen was loaded in the capture tank (Fig. 4) which was then closed and a vacuum drawn. Once the vacuum level reached 100 mBar, the gun was pressurised to the desired pressure and then fired. The sabot/projectile system was accelerated along the 5 m long smooth bore barrel. At the end of the barrel, a ''sabot-stripper" catches the sabot, allowing the projectile to fly further onto the target. A chronograph (Skanar Mk10) was used to detect the impact velocity of the projectile and was placed between the gas gun muzzle and the test panel. The capture chamber has 9 polycarbonate windows to externally view the impact event. Two high speed cameras (Phantom v711) were used to record the impact event. One camera was positioned parallel to the testing rig and used to measure the residual velocity V r of the projectile, and the maximum backface deformation of the impacted panel. The second high speed camera was positioned at an offset angle of $30 in the horizontal plane to observe the back face of the panel during testing. An array of LED lights was mounted on the back of the testing rig to provide enough illumination to record the impact event. In order to hold the target, a steel rig (Fig. 3b) was designed, manufactured and mounted into the capture tank (Fig. 4a). The target panel was sandwiched between two steel frames and simply clamped by bolts. The position of

Data reduction and ballistic curves
In order to estimate the ballistic limit of the manufactured composites, the bisection method was used [15]. This method was selected because: (a) Limited number of test specimens were available; (b) The most appropriate method to find V BL based on cost, accuracy and reliability [15].
After defining an interval in which V BL exists, the bisection method requires the test coupons to be impacted at the average of the velocities at which complete perforation did and did not occur. This procedure was repeated for all the impact tests. Although the ballistic limit V BL is considered to be a statistical velocity at which 50% of the rounds fully perforate the tested panels and 50% of the rounds are held by the panel, V BL was then assumed to lie in between the shots at which perforation did and did not occurred.
The ballistic experimental data were fitted using the Jonas-Lambert model [16] using Eq. (1): where V i is the projectile initial velocity, V BL is the ballistic limit velocity, j and p the ballistic Jonas-Lambert parameters. For nondeformable rigid projectiles, the power p is equal to 2 and Eq. (1)   is known as the Recht-Ipson model [17]. In the present study, the projectiles were not observed to deform during the impact with the composite panels, hence p was assumed equal to 2, for all curves. The parameter j was assumed between 0.85 and 1 to fit the experimental results.

Results and discussions
The tensile properties of dry yarns and microbraids are shown in Fig. 5. It can be observed that the mechanical properties of microbraids depend not only on those of the constituent materials, but also on the braid architecture. The smaller the braid angle a, the higher the stiffness after the jamming point (i.e. the point after which the bias yarns are locked, and the braid angle and braid diameter will not change further) and tenacity. It was also observed that the strain to failure increased with increasing braid angle. Not presented in [7] was the tensile behaviour of dry microbraids having a core of unidirectional fibres of the same material ( Fig. 5e and f). Under tensile loading, these types of microbraids showed a stiffer initial response with respect to coreless microbraids, for both materials. The initial plateau before the jamming point almost disappeared whilst their stiffness decreased with increasing a. The saw-tooth like behaviour of microbraids with a high braid angle would be associated with ruptures of the inner UD core. Although the core failed first, the bias yarns acted as a sheath containing the snap-back of the fibres from their failure point, hence they were still able to carry some load until final failure occurred. Moreover, the higher the bias angle, the higher the number of fractures of the UD core prior to failure. However, the strain to failure of bXYH samples was lower with respect to the maximum strain noted for diamond and regular microbraids, for both materials and for the same braid angle and braid diameter.
This difference increased with increasing a. These results suggest that the UD core could have hindered the complete reorientation of the off-axis fibres to the loading direction, reducing the overall strain to failure for all the investigated microbraids.
The ballistic curves and V BL for Dyneema Ò and Kevlar Ò mBRPC are shown in Figs. 6 and 7, respectively. It appears from Fig. 6a that there was a substantial deviation in the ballistic performance of different Dyneema Ò mBRPC. The V BL ranged between 205 ± 6 m/s for cDA1 and 317 ± 16 m/s for cDC2. It is interesting to note that cDCH and cDC2 specimens had higher V BL with respect to the V BL of cDUD by as much 4.8% and 19.5%, respectively. All other Dyneema Ò microbraid reinforced composites showed worse   ballistic performance with respect to cDUD. On the other hand, the difference in the ballistic limit amongst the Kevlar Ò mBRPC was narrower (Fig. 6b). The V BL ranged between 242 ± 12 m/s for cKHC to 305 ± 11 m/s for cKB2 with four types of mBRPC (cKA1, cKA2, cKB2 and cKC2) having higher ballistic limit with respect to the V BL calculated for cKUD. Among all Kevlar Ò based composites, cKB2 showed the highest V BL which was 18% higher with respect to its baseline.
It is possible to appreciate from Fig. 7 that the ballistic limit of Dyneema Ò -based mBRPC slightly increased with increasing braid angle amongst the composites reinforced with 8 yarn and 16 yarn microbraids, respectively. However, no clear trends can be identified for Kevlar Ò cKY1 and cKY2. Although the higher V BL of cDC2 could be attributed to its higher areal density (2.49%) with respect to the A d of cDUD specimens, the ballistic properties of type cDCH can be stated to be superior with respect to the UD counterpart. This is because of a similar V BL , but smaller A d . Furthermore, the ballistic properties of cKA2 and cKB2 samples can be also stated to be superior with respect to the properties of the cKUD samples. This is due to a higher ballistic limit and smaller areal density observed for the former composites with respect to the reference laminate (Fig. 8). However, it is not possible to establish whether the specimens with similar or lower V BL , but lower areal density with respect to the cXUD samples have better ballistic performance with respect to the UD counterparts because the V BL vs. A d or V f trends are unknown.

A note on the Cunniff's scaling law
Cunniff [18] mathematically demonstrated that the best fibres to be used in textiles for ballistic related applications are those having high toughness, high Young's modulus and low density. He proposed a two dimensionless parameter equation which relates the characteristics of the fibres, armour and threat with the most important design parameter for armours, i.e. V BL : is the product of the specific toughness and the longitudinal wave speed in the fibre; A d the areal density of the armour system; A p and m p the presented area and mass of the projectile, respectively; r; e; q and E the strength, strain to failure, density and Young's modulus of the fibre, respectively. Although Eq.
(2) provides a first approximation for the selection of the reinforcing material in an armour, it could only apply to linear-elastic fibres, it does not implicitly take into account compressive stresses, strain-rate and temperature effects, plasticity and non-linear shear response, and also on the shape and material of the projectile. The relation 0:5r=e used to calculate the toughness of linear elastic materials could not be applied to calculate the toughness of materials having a non-linear tensile behaviour. However, since the toughness can be defined as the area under the stress vs. strain curve, it is possible to rearrange Eq. (3) as: Eq. (4) generalises the Cunniff's parameter U Ã taking into account the effective toughness of the reinforcing material in a composite system. The values of U Ã true calculated from Eq. (4) for dry microbraids are determined by substituting the values of E (from the longitudinal wave speed determined through transverse impact experiments), r (determined by dividing the tensile force by the cross section area of the microbraid A, where A ¼ l=q), and e (obtained from the quasi-static tensile tests on dry microbraids). Fig. 9 shows the Cunniff's parameter with respect to the braid angle for Dyneema Ò and Kevlar Ò yarns and dry microbraids, respectively.
It appears from Fig. 9 that the Cunniff's parameter U Ã overestimates the ballistic performance of an armour system reinforced with microbraids, for both materials. The difference between the values of U Ã and U Ã true increased with increasing braid angle. The trends noted in Fig. 9a for Dyneema Ò -based materials may indicate that the use of unidirectional fibres is recommended in composites for ballistic impact applications. This is associated with the greater ability to absorb and dissipate energy via waves faster with respect to any microbraids, for the same A d ; A p and m p . However, in contrast to what is predicted using Cunniff's U Ã parameter, the ballistic limit of composites reinforced with Dyneema Ò microbraids was similar or higher with respect to V BL of composites reinforced with unidirectional Dyneema Ò fibres. The reason for this mismatch could be attributed to the greater ability of microbraids to deform with respect to the unidirectional counterpart when subjected to external loads. The lower strength and stiffness of the microbraids can be compensated by the higher strain to failure. Other important factors such as yarn-to-yarn friction, geometrical reorientation of the bias yarns and intralaminar strength could have played an important role in the dissipation of the impact energy and must be taken into account in an analytical solution. From Fig. 9b, it appears that the ballistic properties of polymer composites reinforced with Kevlar Ò microbraids must be superior with respect to composites reinforced with unidirectional Kevlar Ò fibres for the same A d ; A p and m p . Moreover, the higher the braid angle, the better the predicted ballistic performance of the composites. However, these results may be misleading due to the fact that they are calculated on the basis of the engineering properties of Kevlar Ò microbraids and do not take into account their effective stress vs. strain behaviour. Calculating the true U Ã true , it appears that the ballistic properties of Kevlar Ò reinforced composites were not significantly affected by the structure and architecture of the reinforcing phase. The trends observed for U Ã true reflected those experimentally noted for cKYZ composites, with the ballistic limit of the Kevlar Ò composites only slightly affected by the type of microbraid used to manufacture the laminates.   From Fig. 10a and d, it is clear that a pyramidal shape deformation developed on the impacted coupons, typical of cross-ply laminates. Delamination started from the impact point and propagated throughout the panels. However, no delamination was seen in the clamped regions of the panels, regardless of the impact velocity. For impact velocities V i 6 V BL , the steel ball slipped through the first layers of the laminate with primary fibres not visually damaged at the impact region. At V i > V BL , although the fibres of the cXUD composites failed at the impact point, it appears that the fibres engaged by the steel ball were not properly loaded but just ''pushed" apart, allowing the threat to slip through the last few layers of the composite laminate. This issue would have detrimentally affected the ballistic properties of the manufactured panels.
The ballistic limit of cDUD composites was much lower with respect to V BL of the commercial Dyneema Ò HB26 laminates for the same areal density. For example, Karthikeyan and Russell [19] noted a ballistic limit of about 410 m/s for HB26 laminates  having an areal density of 2.6 kg/m 2 impacted by steel spheres of diameter 12.7 mm and mass 8.3 g. Moreover, the V BL would have been even higher if smaller and lighter threats were employed in these tests. Although the grade of fibres in cDUD and HB26 laminates are the same, the stacking sequence, thickness of the prepregs, fibre volume fraction, resin type and the possible differences in the consolidation profile could have contributed to the large mismatch between the ballistic performances of the two systems.
No comparison could be made with UD cross-plied Kevlar Ò 49 laminates as no ballistic data could be found in the open literature. Fig. 11 shows a series of snapshots of cDY2 composites impacted at velocities between 232 and 237 m/s. Within the first 100 ls after impact, the deformation mechanisms occurring on Dyneema Ò mBRPC impacted at approximately the same V i were similar amongst the different samples of the same material. The damage created by the impactor was very localised around the region in which the steel ball struck. Delamination was noted around the impact site, which did not propagate to the edges of the panels, regardless of the impact speed and reinforcing microbraid type. Photographs of the front and back faces of cDC2 laminates for penetrating and non penetrating projectiles are shown in Fig. 12.

Dyneema Ò mBRPC
Regardless of V i , the steel ball pushed the microbraids into the panel ( Fig. 12a and b). Only the primary microbraids were drawn in, whilst the secondary microbraids apparently remained unaffected by the impact event. This observation suggests a poor load transfer from the primary microbraids to the adjacent microbraids.
Inspecting the panels impacted at V i 6 V BL , it was not possible to visually observe any failed microbraids, which were simply pushed into the laminate by the impactor. At V i > V BL , although the microbraids directly in contact with the steel ball failed, microbraids on the back face of the laminate were not visually damaged by the impactor, which just slipped through the last layers. Microbraids on the outermost layer disbonded from the penultimate layer and the extent of damage increased with increasing V i .
Amongst the tested Dyneema Ò mBRPC, only type cDC2 showed an out-of-plane displacement greater than 10 mm. Fig. 13 shows the out-of-plane displacement history for an impact just below ballistic limit. The error associated with each point is ±0.69 mm, i.e. the length of four pixels of the recorded high speed videos.
From the results of this investigation, it appears that the resin system constrained the full deformation of the microbraids. A high laminate shear and compressive strength may have hindered the in-plane load transfer from microbraid to microbraid and the out-of-plane deformation. Further research needs to investigate the effect of different resin systems and curing parameters on the ballistic response of Dyneema Ò -based mBRPC.

Kevlar Ò mBRPC
Fig. 14 presents high speed video snapshots showing the impact response of cKY1 mBRPC impacted at V i ranging between 273 and 288 m/s. It can be seen from these snapshots that the impactor created a pyramid-like deformation which grew throughout the unclamped part of the panels. The deformation propagating on the panels reinforced with microbraids having the highest braid  angle was slightly more visible than the deformation growing on the panels reinforced with stiffer microbraids.
A visual post-mortem inspection of the panels revealed that the Kevlar Ò mBRPC impacted at velocity just below V BL did not experience severe damage on their front face (Fig. 15a). However, the primary microbraids in composites cKA1 and cKC1 behaved differently upon impact, with the former being pushed into the panel while the latter being shear cut. It is not possible to attribute this phenomenon only to the reinforcing architecture since the difference in the striking velocity, although small, could have influenced the response of the composite systems. Specimen cKA1 severely delaminated even on the clamped region of the panel with multiple ply splitting occurring. On the other hand, although a clear delamination along the edge of the specimen cKC1 was noted, only one major intralaminar failure was observed.
From Fig. 15c, the drawing in of the edges of the panel was highly visible. As the panel was deforming out-of-plane, the edges of the panel were drawn inward toward the impact point. Contrary to what was seen in impacted Dyneema Ò mBRPC, in which only the primary microbraids moved towards the centre of the panel, the impact event promoted a bigger portion of the panel edges to move in, with a maximum displacement of the outer layer toward the centre of the panel of 8 mm. Although this effect was clearly visible in cKA1 samples impacted at V i 6 V BL , it was less pronounced in composites reinforced with either high a or regular microbraids, and almost completely disappeared for V i > V BL , regardless of the reinforcing architecture.
At V i > V BL , the Kevlar Ò mBRPC showed similar behaviour, regardless of the reinforcing architecture. The primary microbraids on the front face of the panels always failed upon impact,       Fig. 16 presents the maximum out-of-plane displacement history for different cKYZ mBRPC impacted just below the ballistic limit. From these graphs, it appears that the maximum out-ofplane displacement depended on the reinforcing architecture of the microbraid. The smaller the braid angle, the lower the back face deformation of the panel upon impact. cKUD showed the lowest bulge deformation, which reached its maximum value of 16 mm 287 ls after impact. It also appears that the composites reinforced with high braid angle microbraids had a faster deformation response, with an out-of-plane displacement for sample cKC1 reaching 26 mm after 77 ls. Composites reinforced with 8 yarn diamond microbraids showed a higher extent of deformation with respect to composites reinforced with 16 yarn regular microbraids. Moreover, cKUD and cKY2 samples showed a significant springback effect, a phenomenon less pronounced in composites reinforced with finer microbraids.

Hybrid cXYZ concepts
The entry and the exit faces of cDCH and cKCH mBRPC impacted at velocities below and above V BL are shown in Figs. 17 and 18, respectively.
For both materials impacted at V i 6 V BL , the steel ball significantly indented the laminates at the impact point. At V i > V BL , whilst the Kevlar Ò microbraids were cut by the impactor, the Dyneema Ò microbraids fractured ends appeared to be melted. Although little energy was dissipated by membrane deformation and delamination, the ballistic limit of cDCH was greater with respect to V BL of the counterpart made with UD fibres. Not only was the V BL higher, but also the A d was lower with respect to the baseline. Looking closely at the penetration boundaries on the panels impacted above ballistic limit ( Fig. 17b and d), the primary microbraids of the first four layers of the panel appear to have been properly engaged by the steel ball. This phenomenon, as well as friction between fibres and microbraids, could have helped to achieve a greater deceleration of the impactor.
The cXmix concepts arise from the logic of exploiting the potential of unidirectional fibres on the front face of the panel and gradually increasing the strain to failure of the reinforcing layers by using microbraids having higher braid angle though the thickness. In this way, the front face of the panel, which, as previously discussed, should have better ballistic performances with respect to any microbraid reinforced system, would slow down the projectile and allow the highly extensible microbraids to catch the projectile. Fig. 19 and 20 show the front and back face of cXmix concepts impacted above and below V BL , respectively. From these images, little membrane deformation was observed, regardless of the constituent reinforcing material. The calculated V BL for cDmix and cKmix was 23% and 7% lower with respect to their respective baselines. The reason for this can be attributed to the excessive slippage of the projectile through the layers of the laminate.
The bulge deformation of cDCH and cDmix was less than 10 mm and hence not detected. The out-of plane history of cKCH and cKmix is shown in Fig. 21.

Conclusion
In the present investigation, the ballistic response of different microbraid reinforced polymer composites has been experimentally assessed through a series of impact tests. The results showed that, on a weight-to-weight basis, the ballistic limit of certain types of mBRPC were superior with respect to those noted for cross-ply laminates made of unidirectional fibres and manufactured using the same technique.
Dyneema Ò based composites showed very little out-of plane displacement and a poor in-plane load transfer. Very little intralaminar damage was seen around the impact site and it did not propagate up to the edges of the specimens. In the majority of the coupons impacted below V BL , the projectile slipped through and did not engage the first layers of microbraids, and little or no damage observed. At impact velocities above V BL , the microbraids on the front face of the panels directly in contact with the projectile were cut via shear mechanism. Thermal damage was also observed on the front face, especially in core-filled microbraids. The higher the impact speed, the more localised the damage area around the impact site. Among the eight investigated concepts, only two types of Dyneema Ò mBRPC showed a higher ballistic limit, as much as 19.5%, with respect to the V BL of the reference composite made of UD fibres.
Composites reinforced with Kevlar Ò microbraids showed a greater out-of-plane deformation with respect to composites manufactured using Kevlar Ò UD fibres of the same grade, regardless of the microbraid type and impact speed. Delamination propagated from the impact point throughout the panel. At impact velocities below the ballistic limit, little or no fibre damage was visually seen in the impacted panels. On the other hand, fibres and microbraids were shear cut by the steel ball on the front face of the impact coupons and primary microbraids pulled-out on the back face. Amongst the eight investigated concepts, four Kevlar Ò mBRPC had superior V BL , by as much 17.5%, with respect to the reference baseline.
Hybrid composites made using unidirectional fibres and microbraids of different braid angles showed similar or worse ballistic limit with respect to the reference benchmark, for both investigated materials.
Regardless of the reinforcing material and impact velocity, slippage of the projectile through the layers of the composite panels was noted in all tests. This phenomenon could have significantly reduced the ballistic properties of the manufactured composites. Modifications and optimisation of the laminate architecture, curing parameters and resin system are needed to fully exploit the potential of the microbraids in high performance composites. Further research needs to examine the properties of mBRPC under blast loadings and large mass -lower velocity impacts (e.g. bird strike), which could excite larger portions of the microbraids.
Dyneema and DuPont Ò are acknowledged for the provision of Dyneema Ò and Kevlar Ò yarns, respectively.