Projectile strength effects on the ballistic impact response of soft armor targets

Abstract

Upon impact with a target panel, a portion of a projectile's striking kinetic energy is dissipated via heat loss or deformation. Typical ballistic performance determination standards require strict projectile hardness values of Rc 29 ± 2 for consistency and repeatability, but it is of interest to examine if these required hardness values give a lower bound where the ballistic performance determination is independent of the strength of the projectile. In this study, a large range of yield strengths of metallic right-circular cylinders were used to test the effects on the ballistic response of a multi-ply soft body armor. The results show that with an increase in projectile yield strength, the ballistic limit velocity decreases. This degradation in ballistic performance of the soft armor target levels off at higher yield strengths to about 75% of the expected ballistic performance for Rc 29, indicating that there may be a minimum projectile strength after which the influence of strength is no longer significant. The degree of deformation of projectiles during impact is related to the striking velocity and the off-axis failure of the soft armor target material.

Keywords

ballistic performance soft armor energy dissipation yield strength projectile deformation

The significance and importance of soft body armors for personnel protection has only increased multiple-fold over the decades since their first conception. The impact and energy absorption mechanisms during the ballistic penetration of soft armor systems have long been a subject of extensive studies. The exact underlying physics behind the impact and failure phenomena is still not fully understood, and still largely depends on empirical data from extensive ballistic tests. One of the earlier analytical models was proposed by Recht and Ipson,¹ who formulated a model to predict the ballistic limit velocity of a monolithic target panel. The set of equations presented first modeled the perforation process as an inelastic rigid body impact of the impacting projectile cylinder and the target shear plug. Using the conservation of linear momentum and energy, the full energy balance equation of a cylinder impacting a target panel is given by

\frac{1}{2} m_{p} V_{s}^{2} = \frac{1}{2} (m_{p} + m_{f}) V_{r}^{2} + E_{s} + E_{i}

(1)

where m_p is the projectile mass, m_f is the assumed fabric plug mass, V_s is the projectile striking velocity, and V_r is the residual velocity after perforation. E_s is the work done during formation of a target shear plug. In their formulation, E_i is the energy dissipated via projectile deformation and heat generation when the cylindrical projectile impacts a hypothetical free-standing target shear plug. Recht and Ipson expressed this via the equation¹

E_{i} = (\frac{m_{f}}{m_{p} + m_{f}}) \cdot \frac{1}{2} m_{p} V_{s}^{2}

(2)

During ballistic impact, two possible damage and deformation scenarios may occur: loss of projectile mass due to erosion, or an increase in the presented area due to mushrooming. The former scenario is relatively straightforward, as projectiles have been known to experience mass loss from erosion when impacting a target at high velocities, especially for hard targets.² The projectile gradually loses both mass and velocity during the perforation process and, therefore, some portion of kinetic energy. The latter scenario results in ‘mushrooming’ of the impacted end. This mechanism dissipates striking kinetic energy via plastic work done by deforming the impact end of the cylinder. Consequently, this increased presented area due to mushrooming further results in a higher ballistic limit of the remaining plies in the target by involving more material during perforation. A more detailed energy balance equation that included these other mechanisms was further expressed by Corran et al.³ for projectile impact onto a plate. In their analysis, the projectile's striking kinetic energy was partitioned into target elastic energy, target plastic energy due to permanent plate bending, work done due to shear plug formation, and projectile mushrooming at the impact end.

For soft armor targets, Cunniff⁴ proposed a semi-empirical energy balance that may be similarly expressed as per Equation (1) by rearranging the equations provided therein

\begin{matrix} \frac{1}{2} m_{p} V_{s}^{2} = \frac{1}{2} (m_{p} + X_{2} m_{f}) V_{r}^{2} \\ + \frac{1}{2} m_{p} V_{c}^{2} exp [- X_{3} (\frac{V_{s}}{V_{c}} - 1) x_{4}] \end{matrix}

(3)

V_{c} = X_{5} exp [X_{6} η^{X_{7}}], η = m_{f} / m_{p} = A_{d} A_{p} / m_{p}

(4)

In Equations (3) and (4), X₂ to X₇ are regression coefficients and V_c is the critical ballistic limit velocity. The dimensionless parameter η is the mass ratio of the target plug to the projectile, A_d is the areal density of the target, and A_p is the projectile's presented area. The second exponential term in Equation (3) implicitly includes the work terms E_i and E_s from Equation (1), and possibly other mechanisms that may exist. For the particular case of soft armor targets, the work done E_s due to the formation of a hypothetical shear plug has been shown to be somewhat independent of striking velocity,⁵ and is almost purely a function of projectile diameter, target thickness, and target through-thickness shear strength. Similar effects of projectile plastic deformation on the ballistic performance of soft armor targets have not been of particular focus in the existing literature.

Experimentally, the interactions between these mechanisms may be complex and difficult to isolate, and for this reason soft armor target impact studies have been largely focused on the strength and material properties of the target.^4,6,7 However, the significance of the projectile strength on the ballistic performance should not be ignored, since commercial bullets (such as full metal jacket (FMJ) rounds or semi-jacketed hollow point (SJHP) rounds) are typically made of softer metals and thus deform easily upon ballistic impact (Figure 1). Ballistic limit determination tests are often based on the specified standard hardness of Rc 27–31,^8–10 and while these hardness values are higher than those of commercial bullets, higher impact velocities will also result in projectile deformation.

Figure 1.

Post-impact 9 mm full metal jacket (a) and .44 Magnum semi-jacketed hollow point (b) rounds after impacting a soft armor ballistic vest. Both show extreme deformation due to the low strengths of the lead core and copper jacket.

An earlier study that reflects these mechanisms⁷ was performed by Cunniff using 2-, 4-, and 16-grain right-circular cylinders (RCCs) and, consequently, different target/projectile areal density ratios, on a Pyrex glass/Kevlar® KM2 fabric hybrid target system. The results showed that at lower areal density ratios, that is, larger projectile areal density for the same target system, the full Kevlar® KM2 fabric system outperformed the Pyrex/Kevlar® KM2 hybrid panel (Figure 2). At higher areal density ratios, the Pyrex/Kevlar® KM2 hybrid outperformed the full KM2 fabric system. One mechanism may be due to the hard/brittle Pyrex® layer in a hybrid system that helps resist the projectile at the strike-face more efficiently at higher impact velocities (since higher areal density ratios lead to higher ballistic limit velocities). Cunniff attributed this superior performance of the hybrid to the deformation of the smaller 2- and 4-grain projectiles during impact, whereas the larger 16-grain projectiles do not typically deform within the range of striking velocities tested near the respective ballistic limit. Although further examination of the recovered projectiles was not performed, projectile erosion and plastic mushrooming scenarios nonetheless remain distinct possibilities as well, which is ultimately related to the strength of the impacting projectile and the interaction with the target panel.

Figure 2.

Ballistic limits of full Kevlar® KM2 fabric and Pyrex®/Kevlar® KM2 hybrid, with the latter exhibiting superior ballistic performance when projectiles are small.⁷

In recent studies directly examining the effects of projectile strength, Ćwik et al.^11,12 impacted Dyneema® HB26 and Spectra® 3124 ultra-high molecular weight polyethylene (UHMWPE) composites using steel and copper 20-mm fragment-simulating projectiles (FSPs). The masses of these FSPs are 53.1 and 60.25 g respectively, which indicate relatively similar areal density ratios in their study. The ballistic performance of the copper FSPs was reduced in comparison to the steel FSPs, as the ballistic limit velocity of the target panel was much higher for the copper FSP than the steel FSP. Mass losses were insignificant for both projectile materials (up to 4%) even at high striking velocities, but the copper projectiles were observed to deform substantially via mushrooming. The larger effective projectile presented area due to mushrooming resulted in a larger contact area during impact, and thus a larger target area that failed via tearing or melting. In this study, we further investigate the effects of projectile deformation and mass loss on the ballistic performance of these projectiles. Metal projectiles of various materials and strengths are used to impact and perforate Twaron® soft armor targets.

Experimental procedure

Gas gun setup

Projectiles were shot with a single-stage smooth-bore light-gas gun. The target is located approximately 0.4 m from the tip of the barrel to improve accuracy and reduce trajectory instability and tumbling of the RCC projectiles. Alignment was performed to ensure perpendicularity of the target panel to the shot axis. A steel soft-catch safety chamber was placed behind the target panel mount to retrieve any perforated projectiles. Target panels were clamped with L-brackets and 25.4-mm width Neoprene rubber linings (50A Durometer) for added grip, leaving an exposed surface area of 0.254 × 0.254 m² (10 in. × 10 in.). L-brackets were secured using 12 flanged screws equally spaced on all corners and torqued to 2.8 N-m (25 in-lb). Laser diodes were used to measure velocities accurate to within 3.1 m/s (10 ft/s).

Target material

Base Twaron® CT709 balanced plain-weave fabric samples were made from 930 dtex yarns (27 × 27 ends/picks per 25.4 mm). Fabric samples had an areal density of 4.354 kg/m² for 22 plies. Target panels were cut to 0.305 × 0.305 m² (12 in. × 12 in.) size and edge-stitched three times together with a 25.4 mm (1 in.) margin from the edges for easier handling. Panels were kept in a cool, dry area at room temperature for at least 24 hours prior to shooting.

Projectiles

To study the effect of material strength on the deformation of the impacting projectile, RCCs of different materials were chosen based on different hardnesses and densities, with their respective properties given in Table 1. Rockwell C hardness values for O1 steel RCCs were tested and averaged over several measurements using a spot hardness tester on the impact end. To obtain physically meaningful values for comparison, O1 steel Rockwell C hardness values were first converted to Vickers diamond pyramid hardness values.¹³ The hardened O1 steel yield strengths were then calculated from the Vickers hardness H_v via known correlations provided by Pavlina and van Tyne¹⁴ (Equation (5)). The calculated yield strength was compared with existing studies for Rc 60 O1 steel¹⁵ and AISI M2 steel¹⁶ and was shown to be accurate. Equation (5) also gives an approximate yield strength of 755 MPa for standard Rc 29 projectiles

σ_{y} = - 90.7 + 2.876 H_{v}

(5)

Table 1.

Projectile materials and properties used in the study

Material	RCC/total mass [g]	Rockwell	Vickers	σ_y [MPa]
O1 steel	4.48/4.88	HRC 61	720	2200¹⁵
O1 steel	4.48/4.88	HRC 42	412	1100
M2 steel	4.60/5.00	HRC 62	746	2700¹⁶
360 brass	4.83/5.23	HRB 75	N/A	159
7075-T6 Al	1.60/2.00	HRB 72	N/A	427
6061-T6 Al	1.49/1.89	HRB 60	N/A	241

RCC: right-circular cylinder.

RCC pieces had nominal diameters and lengths of 9 mm. RCCs were tumbled with ceramic media for 5 hours prior to testing to reduce the edge sharpness, as a previous study by the authors showed that localized edge geometries demonstrably reduced ballistic performance of the fabric due to off-axis failure.¹⁷ Tumbled projectiles had micro-scale edge radii of curvature of at least 125 µm, where the ballistic limits are less dependent on these micro-scale stress concentrations. Instead of using sabots, which require stripping before impact, light copper gas checks were lightly attached on the non-impact end using petroleum jelly to form a better gas seal for higher gun efficiency and to achieve higher velocities. These gas checks weigh 0.4 g and have been added to the total mass in Table 1.

Shooting procedure

Twelve shots per panel were performed to determine the ballistic limit using the bracketing method, as detailed in NIJ-0101.06,¹⁸ with the first shot targeted at 305 m/s (1000 ft/s). The shot locations were located 25.4 mm (1 in.) from the panel stitching and at least 50.8 mm (2 in.) apart from each other, and as far as possible, shots were located such that the principal yarns did not overlap. For uniformity in testing, pre- and post-test temperatures and humidity levels were also recorded to ensure that testing conditions did not vary significantly.

Results and discussion

Recorded pre- and post-testing temperatures and relative humidity levels were between 17.0 ℃ and 26.0 ℃ and 34% and 49%, respectively. Tests were completed within 5 hours of test commencement. High-speed images were taking using a Shimadzu HyperVision HPV-X2 to ensure normal impact of the projectile without any significant yaw or pitch in the flight trajectory (Figure 3).

Figure 3.

High-speed image sequence of a 6061-T6 right-circular cylinder impacting a fabric target at 339 m/s, with a frame rate of 400 kHz and 200 ns exposure. A brief flash occurs at the time and site of impact (t = 0). Principal yarns in the vertical direction appear to be strained first before a square pyramidal tent propagates from impact site.

Impact flash phenomena

The image sequences revealed a phenomenon whereby a transient flash of light occurred at the time and site of impact, which only occurs very briefly for a maximum duration of about 5 µs. A similar flashing phenomenon was previously observed by Chocron et al.¹⁹ when impacting Dyneema® HB80 laminates with a polyurethane matrix, and recently by Ćwik et al.¹¹ and Yang and Chen²⁰ on Dyneema® SB71 laminates. Chocron et al.¹⁹ and Ćwik et al.¹¹ attribute this to isentropic shock loading of the polyurethane matrix upon impact, and the flash is a result of an ‘autoignition effect’ from the shock, resulting in localized melting of either the UHMWPE fibers or the polyurethane matrix, or both.

The same phenomenon was observed in previous studies^17,21 when firing O1 steel projectiles on 22-ply Twaron® CT709 fabric, indicating that this phenomenon may be projectile-independent, that is, a flash occurs in the fabric under certain conditions regardless of the projectile material. As far as the authors are aware, there are currently no prior reports of similar phenomena occurring for aluminum projectiles impacting aramid fibers.

Ballistic limit results

The outcome of each shot was assigned a value of ‘0’ for partially perforated shots and ‘1’ for completely perforated shots. Perforation of the panel was verified visually during the test, and via post-mortem for confirmation. Two different methods of calculating the V50 ballistic limit, the NIJ-0101.06¹⁸ and MIL-STD-662F²² methods, were compared and averaged (Table 1). NIJ-0101.06 uses a logistical S-curve regression, while MIL-STD-662F uses the arithmetic mean of the lowest complete penetration velocities and highest partial penetration velocities. The full details of the calculation methods are given in their respective references.

In a previous study, Cunniff²³ demonstrated that the V50 ballistic limit velocities of soft armor targets may be collapsed onto a single curve when the V50 velocities are non-dimensionalized with respect to the Cunniff velocity, that is, the cube root of the product of the fiber specific toughness and the longitudinal sound speed within the fiber, given as

Ω^{1 / 3} = (\frac{σ ɛ}{2 ρ} \cdot \sqrt{E / ρ})^{1 / 3}

(6)

where σ is the longitudinal failure strength, ɛ is the longitudinal failure strain, ρ is the fiber density, and E is the longitudinal modulus. The normalized ballistic limits for several different armor materials and constituent fibers were shown to collapse into a single normalized regression curve for Kevlar® 29 (Ω^1/3 = 624).²³ For Twaron® CT2040, the fiber properties are obtained from Mayo and Wetzel²⁴: failure strength σ = 3.3 GPa, failure strain ɛ = 3.3%, density ρ = 1440 kg/m³, and longitudinal modulus E = 90 GPa, giving Ω^1/3 = 668. The V50 velocities for Twaron® CT709 are then predicted using the equation

V_{50} = \sqrt[3]{Ω_{Twaron} / Ω_{Kevlar 29}} \cdot [X_{5} exp (X_{6} η^{X_{7}})]

(7)

X₅, X₆, and X₇ are the regression coefficients given by Cunniff⁴ as 269.32, 2.9068, and 0.7586, respectively. Since the ballistic data is typically based on a standardized projectile Rockwell C hardness of 27–31, these predicted V50 values provide a baseline comparison for the experimentally obtained V50 in our experiments to account for different areal density ratios. The experimental and predicted V50 velocities and the experimental to predicted ratios are also tabulated in Table 2. Experimental data for Twaron® CT709 was plotted along with the regression curve in Figure 4, calculated using Equation (7), as well as with data from previous studies¹⁷ using 7-mm O1 steel RCCs for comparison.

Figure 4.

Plot of Twaron® CT709 V50 against areal density ratio η, with the regression curve calculated using Equation (7) for comparison.

Table 2.

Experimental and predicted ballistic limit velocities and ratios

		V50 [m/s]
Material	100 × η	NIJ	662F	Exp.	Pred.	Ratio
O1 steel, Rc 61	5.68	345	349	347	388	0.895
O1 steel, Rc 42	5.68	337	339	338	388	0.871
M2 steel	5.54	316	316	316	393	0.804
360 brass^a	5.30	411	410	>411	381	1.078
Al 7075-T6	13.85	606	599	603	533	1.131
Al 6061-T6^b	14.65	–	–	>601	548	1.096

Tests were performed for 12 shots up to a maximum of 410 m/s with only one complete penetration. The ballistic limit is assumed to be at or above this velocity.

Tests were performed for only four shots. Maximum possible velocities of 601 m/s were achieved without complete penetration. The ballistic limit is assumed to be at or above this velocity.

The brass and aluminum projectiles lie above the master curve, indicating subpar performance of these projectile materials in comparison to a baseline Rc 29 steel projectile. On the other hand, the stronger projectiles lie at or below the predicted V50, indicating similar or superior performance. By plotting the experimental to predicted V50 ratios from Table 2, the influence of the areal density ratio η on the target performance results is eliminated, and the effects of strength on the ballistic performance may be isolated – these results are plotted in Figure 5. The figure also includes the predicted point for a RCC with hardness Rc 29 (yield strength 755 MPa) and using Equation (7).

Figure 5.

Plot of V50 ratios against the yield strengths of projectiles.

As expected, the general trend of the ballistic limit velocity ratio decreases with an increase in projectile yield strength. Although the somewhat linear data appear to indicate that a hypothetical infinitely high-strength projectile impacting the target would yield a near-zero V50, it is more likely that the influence of projectile strength would level off, closer to what the power-law fit in Figure 5 would suggest. Note that this fit is merely included to visualize trends and should not be taken to be predictive. The effects of inelastic impact deformation are further investigated via post-mortem analysis of the projectiles.

Projectile deformation

Post-impact projectiles were caught with a catch chamber lined with cotton jersey cloth rags and soft rubber sheets to prevent deformation of the projectiles upon impact with the chamber walls. On occasion, partially penetrated projectiles were trapped within the fabric target layers, and subsequent shots may impact these trapped projectiles, resulting in extreme deformation and damage. These projectiles were not considered for post-mortem analysis. Post-impact diameters, lengths, and masses were measured to quantify the amount of deformation or damage occurring to the projectiles during impact (Table 3). Note that the average values for each material type in Table 3 are taken across a range of impact velocities.

Table 3.

Post-impact measurements of right-circular cylinders

		Post-impact diam.		Post-impact length		Post-impact mass
Material	No. of points	Meas. [mm]	Ave. % change	Meas. [mm]	Ave. % change	Meas. [g]	Ave. % change
O1 steel, Rc 61	13	8.99 ± 0.01	–0.11	8.99 ± 0.01	–0.11	4.48 ± 0.01	0.00
O1 steel, Rc 42	10	9.00 ± 0.01	0.00	9.00 ± 0.01	0.00	4.48 ± 0.01	0.00
M2 steel	12	9.00 ± 0.00	0.00	8.99 ± 0.02	–0.11	4.59 ± 0.01	–0.22
360 brass	12	9.05 ± 0.09	0.56	9.00 ± 0.02	0.00	4.82 ± 0.01	–0.21
7075-T6 Al	9	9.05 ± 0.04	0.56	9.00 ± 0.02	0.00	1.60 ± 0.00	0.00
6061-T6 Al	4	9.81 ± 0.42	9.00	8.85 ± 0.20	–1.67	1.49 ± 0.00	0.00

From Table 3, the mass loss during ballistic impact is practically negligible, implying that the projectiles, even the high hardness Rc 61 RCCs, do not fail in a brittle fashion, but rather they deform in a ductile fashion if at all. The post-impact diameters are noticeably larger for softer materials, that is, brass, 7075-T6, and 6061-T6. An explicit analytical expression for the projectile mushrooming energy was previously given by Johnson²⁵ and is included in Corran et al.'s analysis³ for projectile impact on ductile steel plates

\begin{matrix} E_{m} = \frac{2 σ_{y} A_{p} H}{9 (\sqrt{A_{p}, d / A_{p}} - 1)} \\ \cdot {3 (\sqrt{\frac{A_{p, d}}{A_{p}}}) 3 \ln [\sqrt{\frac{A_{p, d}}{A_{p}}} - (\sqrt{\frac{A_{p, d}}{A_{p}}}) 3 + 1]} \end{matrix}

(8)

In Equation (8), A_p,d is the projectile's mushroomed presented area and H is the length of the deformed section. While A_p and A_p,d can be easily calculated from the measurements in Table 3, the difficulty lies in measuring and determining the deformed section length H, as the perforation process typically results in non-ideally axisymmetric impact for these soft armor targets. Equations (2) and (8) together suggest that the trends of the plastic work done is related to the striking kinetic energy, and therefore the post-impact percentage change in A_p for each shot and each material are plotted against the striking kinetic energy in Figures 6(a)–(d). Since none of the steel projectiles exhibited significant deformation, only the Rc 42 O1 steel RCC data is plotted for comparison.

Figure 6.

Percentage change in the projectile presented area A_p against the striking kinetic energy for (a) Rc 42 O1 steel, (b) 360 brass, (c) 7075-T6, and (d) 6061-T6.

Slight negative changes in A_p are due to deviations in the measured diameters. The soft metals exhibited much larger degrees of deformation as the inelastic impact energy increased. For the aluminum RCCs, the areal density of the fabric target is relatively higher compared to their respective projectile masses, which resulted in larger degrees of deformation. However, 7075-T6 displayed less yielding compared to 6061-T6 due to the higher strength of the former. Post-impact projectiles were examined for micro-scale damage or deformation via scanning electron microscopy (Nova NanoSEM 200, Thermo Fisher Scientific, USA) in Figures 7–9.

Figure 7.

Post-impact micrographs of a M2 tool steel right-circular cylinder projectile shot at (a) 306 m/s (sub-V50), striking KE 234 J, and (b) 387 m/s (above V50), striking KE 374 J. Negligible to no deformation is observed at either velocity.

Figure 8.

Post-impact micrograph of a 360 brass right-circular cylinder projectile shot at 298 m/s (sub-V50), striking KE 232 J. Slight mushrooming deformation is observed at the impact end. Yarn imprints may be observed as well.

Figure 9.

Post-impact micrographs of 7075-T6 aluminum right-circular cylinder projectiles shot at (a) 387 m/s (sub-V50), striking KE 150 J, and (b) 620 m/s (above V50, complete penetration), striking KE 384 J. Deformation may be observed for both velocities, although larger degrees of mushrooming are observed at velocities above V50.

Micrographs of steel projectiles showed minimal damage/deformation at the corners, which is reflected in the post-impact measurements in Figure 7. For the brass and 7075 projectiles (Figures 8 and 9, respectively), micro-scale plastic deformation was observed even though the 7075 projectiles are much stronger than the brass projectiles. The micro-scale deformation for these softer projectiles (in comparison with the much stronger steel) is related to the striking kinetic energies, as suggested by Equation (2). In general, where the projectile hardness exceeds the standardized values of Rc 29 ± 2, the projectiles do not appear to exhibit any large-scale deformation, even at velocities higher than the ballistic limit. For softer projectiles, the impact ends generally deform via mushrooming. The deformation is obvious when examining the post-impact 6061-T6 projectiles, where the impact end severely mushroomed out and the fabric weave pattern became imprinted (Figure 10). Although not to such a severe degree as the 6061-T6 projectiles, the brass projectiles showed similar weave-pattern imprinting and mushrooming. This is likely due to the range of striking kinetic energies for brass being relatively low compared to the 6061-T6, as seen in Figure 6.

Discussion of energy dissipation due to damage or deformation

Target post-mortem analysis

It was also observed that the degree of deformation appears to be somewhat related to the impact velocity and the outcome of the shot (i.e. partial or complete perforation). These data points are shown in Figures 6(b) and (c), where some of the high-impact kinetic energies yielded minimal deformation. To explain this behavior, post-mortem inspection was also performed on the impacted targets. An examination of the fabric target's failure modes using prior methods for all shots reveals that the constituent plies of woven fabric targets typically respond in three broad ways: via yarn rupture, nosing-through (also known as ‘windowing’), or no failure,^17,26 as shown in Figure 11. At high velocities with respect to the individual ply, the yarns fail and rupture locally without significant yarn pull-out (Figure 11(a)). This type of failure mode typically occurs for the frontal few plies closer to the strike-face. When the velocity is sufficiently low, the yarns do not fail, but slip through the weave structure instead. A certain degree of yarn pull-out is often observed along with the characteristic diamond-shaped region of fabric strain (Figure 11(b)). This defeat mode typically occurs in the middle of the pack after the regime of yarn rupture, and for shots that completely perforate at or near the critical velocity, this form of ply failure may be found at the rear of the pack. When the projectile is finally too slow to initiate sufficient yarn pull-out to slip through the weave structure, only minor dents and deformation are observed (Figure 11(c)). The degree of deformation is difficult to quantify, but in this mode, it is obvious that the projectile does not perforate the fabric. Since this inspection method is purely visual, it should be noted that large scatter in determining actual failures modes is to be expected.

Figure 10.

Photographs of (a) top and (b) elevated side profiles of post-impact 6061-T6 aluminum right-circular cylinders (RCCs). The striking velocities of RCCs were (left to right) 364, 490, 561, and 601 m/s, respectively. Progressively severe degrees of mushrooming deformation are shown with increasing striking velocities.

Figure 11.

Broad categories of failure modes as observed in post-mortem: (a) localized yarn rupture; (b) windowing/nosing-through; (c) no failure.

Figure 12 shows that, generally speaking, the percentage of ruptured plies increases with V_s, with practically all the plies failing via rupture (i.e. off-axis modes) at velocities past the V50. The 6061-T6 aluminum was not included in Figure 12 as the V50 velocity is expected to be much higher, since practically no yarn rupture was observed even at the maximum velocity of 601 m/s. The failure mode trends suggest that, at high velocities, the localized failure of the contacted target material results in less projectile edge damage/deformation. Near or below V50 velocity regimes, the yarns of the individual plies are more likely to survive the impact for a longer time. This sustained contact time of the target and the projectile results in larger degrees of deformation of the projectile before the target material strains to failure. In cases where localized off-axis failure occurs and the target is defeated, energy dissipation via projectile deformation is lower compared to cases where the projectile is halted by the fabric. For example, a brass RCC with V_s = 391 m/s completely perforated the 22-ply Twaron® panel and ruptured approximately 91% of the plies, but exhibited no significant post-impact change in diameter; for another partially perforated shot with V_s = 401 m/s, the brass RCC ruptured 41% of the plies but had a post-impact diameter of 9.27 mm (3% change). Presently, the dependence of the degree of yarn rupture on the projectile strength is inconclusive with the current set of post-mortem data and should be examined via in situ methods.

Figure 12.

Plot of the percentage of ruptured plies against normalized striking velocities.

Conclusions

The material yield strengths of RCC projectiles were varied to investigate the role of projectile deformation as an energy-dissipation mechanism during soft armor target impact. Target thicknesses were kept constant to isolate and exclude the effects of through-thickness shearing. The ballistic performance of the target panel generally decreases with an increase in projectile yield strength, although the marginal improvement in projectile performance is observed to diminish at higher yield strengths. The measured post-impact degrees of deformation of the RCCs correlate well with the inelastic impact energy dissipation using Recht and Ipson's formulation, especially for lower strength projectiles. At velocities near V50, localized failure of the target material resulted in less severe deformation of the impinging projectiles compared to similar velocities where the projectile was stopped. This effect is more pronounced for softer projectile materials, where yielding is more likely to occur at these ballistic velocities.

Footnotes

Acknowledgements

The authors would like to thank the US Army, P.M. Soldier Protection and Individual Equipment, Technical Management Directorate for their support. The authors would also like to thank Joan Goetz of Purdue University’s Department of Consumer Science for her help with stitching and sewing the fabric target panels.

Declaration of conflicting interests

The authors declared no potential conflicts of interest with respect to the research, authorship, and/or publication of this article.

Funding

The authors received no financial support for the research, authorship, and/or publication of this article.

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