Design and development of ceramic/metal/UHMWPE hybrid ballistic panels for ground combat vehicle applications: Experimental testing and numerical modeling

Abstract

Metallic panels are often used for vehicle protection, designed to withstand a range of projectiles, from light firearms to armor-piercing (AP) rounds. However, these panels significantly increase the vehicle’s weight, potentially degrading performance and payload capacity. The armored vehicle market demands lighter, more fuel-efficient vehicles. This study focuses on optimizing a reference ballistic steel panel (RBP) by integrating ultra-high molecular weight polyethylene (UHMWPE) and boron carbide (B₄C) ceramics to meet STANAG 4569 Level 3 standards while reducing areal density (AD) for ground combat vehicle (GCV) applications. Material characterization was performed on reference materials using a universal testing machine (UTM) and a Split-Hopkinson pressure bar (SHPB) with 3D digital image correlation (DIC) to evaluate performance at different temperatures, strain rates, and failure modes under varying stress triaxiality. The identified material parameters were used to develop finite element (FE) models, which were validated and then employed to optimize several panel designs. Three promising configurations identified through FE modeling include (i) BP1 - (BS600/BS550/UHMWPE), (ii) BP2 - (B₄C/BS600/BS550), and (iii) BP3 - (B₄C/BS600/BS550/UHMWPE). These configurations were manufactured and tested against AP rounds using a gas gun projectile launcher. As a result, two panels were developed with reduced areal densities, 26% and 37% lighter than the reference ballistic steel panels, while maintaining superior ballistic protection. The single-hit impact test data from this study will aid in the design and manufacture of a lightweight, highly protected GCV.

Keywords

ballistic panel projectile vehicle armor high strain rate numerical simulation

Introduction

In recent years, the use of armored vehicles worldwide has increased due to several factors, including ongoing global geopolitical instability, increased automation, availability of advanced defense equipment, and lightweight materials Li et al. (2025), Giurgiu et al. (2023). In a recent article (The Business Research Company (2026)), it was reported that the armored vehicles market is expected to grow at a 5% compound annual rate from 2025 to 2033. These vehicles are equipped with ballistic protective panels designed to shield passengers from projectiles for military, security, and civilian purposes Tsirogiannis et al. (2024). The primary goal of military armored vehicles is to stop armor-piercing (AP) projectiles while simultaneously reducing weight Acar et al. (2024), Ozer et al. (2025). The ballistic panels are manufactured in accordance with various standards, including the National Institute of Justice (NIJ) 0101.07 Mukasey et al. (2008) and STANAG 4569 Burian et al. (2019), which specify their performance requirements. The STANAG 4569 standard is commonly used in the design and manufacture of armored vehicles Vilis et al. (2025). Particularly, STANAG Level 3 and 4 are the recommended standards for any ground combat vehicle applications Tsirogiannis et al. (2024).

STANAG 4569 Level 3 requires effective protection against 7.62 x 51 mm AP with tungsten carbide core projectile (AP-WC) at 930 m/s (Burian et al. (2019)) and 7.62 x 54 API B32 (API-B32) projectile at 854 m/s Kilic and Ekici (2013). In the past, several researchers have developed ballistic panel configurations to withstand AP-WC projectiles with areal densities ranging from 140 to 177 kg/m² Klement et al. (2008), ArcelorMittal (2024) and to withstand API-B32 projectiles with areal densities ranging from 75 to 125 kg/m² Strasburger (2009), Silva et al. (2014), Harris et al. (2017), SWEBOR ARMOR™ (2026).

Several researchers reported ballistic panel configurations designed to withstand AP-WC and API-B32 projectiles, as summarized here. Mars® steels from ArcelorMittal (2024) reported that the multi-layered ballistic panel configuration [(BS600)_{9 mm}/(AG)_{15 mm}/(BS380)_{13.5 mm}] with the areal density of 177 kg/m² could withstand the AP-WC projectile. TRICERA Inc (2026) reported that the ballistic panel manufactured with [(synergic ceramic metal composite)_{23 mm}/(Armox 500)_{7.5 mm}] with the overall thickness of 31 mm and areal density of 140 kg/m², could withstand the AP-WC projectile. Harris et al. (2017) developed the armor configuration [(Aramid)_{0.6 mm}/(silicon carbide (SiC))_{8 mm}/(epoxy)_{0.65 mm}/(Aramid)_{2 mm}/(E-glass fiber)_{25 mm}] with the total thickness of 36.25 mm and the areal density of 93 kg/m². This armor demonstrated multi-hit capability against the API-B32 projectile. However, the study recommended additional testing to ensure strong adhesive bonding between the ceramic and the composite. SWEBOR ARMOR™ (2026) reported that replacing the solid 16 mm (500 Brinell type-armor steel) with a combination of a 6.5 mm (500 Brinell type-armor steel) layer and a 4 mm (perforated 600 Brinell type-armor steel) layer can meet the API-B32 standard. This change reduces the areal density from 125 kg/m² to 75 kg/m², a 40% reduction. Strasburger (2009) manufactured panels with alumina ceramic strike-face materials, thicknesses ranging from 1.3 mm to 8 mm; soda-lime glass (30 mm) as the intermediate layer; and polycarbonate (4 mm) at the back face. The areal density of these panels varied from 85 kg/m² at 35.3 mm thickness to 112 kg/m² at 42 mm. Tests were conducted at the impact velocity of 850 ± 15 m/s using the 7.62 × 51 mm AP-hardened steel-core projectile, similar to API-B32. The results showed that impact resistance increases with ceramic thickness up to 4 mm, beyond which further thickness yields little to no additional benefit. Klement et al. (2008) manufactured a 33 mm thick [(Al₂O₃)_{7 mm}/(float glass)_{24 mm}/(polycarbonate)_{2 mm}] ballistic panel that corresponds to an areal density of 90 kg/m², designed to withstand the API-B32 projectile. However, a significantly thicker ballistic panel, 55 mm thick with an areal density of 150 kg/m², was manufactured using the same material combinations [Aluminum oxide (Al₂O₃)_{7 mm}/(float glass)_{48 mm}/(polycarbonate)_{2 mm}] to withstand an AP-WC projectile. It is important to note that significantly higher thickness and higher areal density are required to stop the AP-WC projectile compared with the API-B32 projectile. Despite being able to withstand STANAG Level 3 protection, these design configurations sacrifice higher area density. The higher areal density can negatively affect the vehicle’s performance, maneuverability, and payload capacity.

Currently, most armored vehicles are made of armored steel Acar et al. (2024), Ozer et al. (2025), Orlov et al. (2023), Ryan et al. (2016) and a combination of steel/ceramics Silva et al. (2014), Khan et al. (2020). To withstand high-caliber AP projectiles, either thicker multiple-steel panels or multi-layered ceramic and steel panels are necessary, which add weight Tsirogiannis et al. (2024), Ozer et al. (2025), Vilis et al. (2025). Thinner steel panels are inadequate against armor-piercing rounds Acar et al. (2024). Therefore, there is a need to explore alternative materials to upgrade armored vehicles worldwide, thereby reducing their overall weight. Combining different multi-layered material configurations, such as ceramics, fiber reinforced polymer (FRP) composites, and metallic alloys, enables the creation of ballistic protection structures that are lighter and possess higher specific energy absorption than those made from traditional metallic materials Tsirogiannis et al. (2024), Vilis et al. (2025), Aziz et al. (2025), Li et al. (2020). Tsirogiannis et al. (2024) reported that advanced FRP composites, particularly aramid and ultra-high molecular weight polyethylene (UHMWPE) based composites, have been hybridized with metal/ceramic structures to reduce the projectile’s perforation capability effectively while minimizing the weight of the ballistic panel in vehicle armor applications.

Generally, the strike face for multi-layered configurations is made of high-hardness materials, such as specialized armored steels or ceramics, to erode AP projectiles upon impact Vilis et al. (2025). Al₂O₃, B₄C, and SiC are the commonly used ceramics for the strike face of the ballistic plates owing to their excellent mechanical properties, particularly their superior stiffness and hardness against the armor-piercing projectiles Khan et al. (2020), Zou et al. (2023). Dresch et al. (2024) compared the impact performance of a monolithic ceramic ballistic plate with a mosaic ceramic ballistic plate. The mosaic ceramic ballistic plate consists of tiles arranged side by side. This tile arrangement provides a longer interaction time with the projectile. As a result, better impact resistance can be achieved, which is an advantage for multi-hit applications compared with a monolithic ceramic ballistic plate. Pai et al. (2022) and Marpaung and Aritonang (2023) reported that advanced fiber composites, such as UHMWPE, aramid, and nylon, can be effectively used as ballistic backings for the spall liner to enhance armor protection. However, selecting the appropriate materials and optimizing the design can be challenging and time-consuming, often resulting in suboptimal configurations. Moreover, ballistic phenomena are complex and influenced by factors such as high strain levels, strain rate effects, nonlinear behavior, shock waves, and boundary constraints Burian et al. (2019). Although many damage models exist today, incorporating the significant impact of triaxiality on the damage and fracture of metals is crucial for most of them Rickhey and Hong (2022). The fracture mechanics community proposed stress triaxiality and the Lode angle as two stress-state-dependent parameters that affect material fracture behavior Anderson et al. (2017), Li and Jing (2022).

Stress triaxiality is defined as the ratio of hydrostatic stress to von Mises equivalent stress, providing insight into stress distribution and failure conditions in ductile materials. The fracture process in ductile metals can be examined through experimental measurements of stress triaxiality using various specimens Benzerga and Leblond (2010). In our previous study Sartor et al. (2022), seven different Ti6Al4V specimen configurations were manufactured, and then tensile tests were carried out to cover a range of stress triaxiality levels from −0.28 to 0.90. Additionally, the parameters for both elastoplastic and damage behaviors were determined and employed in FE-based numerical models to predict the mechanical behavior of the tested specimens. The FE modeling yielded satisfactory simulation results for specimens with diverse configurations when subjected to high strain levels and varying triaxialities. These factors, including strain rate and temperature effects, boundary constraints, and stress triaxiality, must be considered in the FE modeling process. There are very few studies that integrate stress triaxiality-informed damage calibration into ballistic FE modeling. This approach is crucial for bridging the knowledge gap and refining panel designs to improve protection while reducing weight and cost.

In this work, mechanical characterization was performed using a UTM and an SHPB to provide input data for FE modeling. Tests assessed material response at different temperatures, strain rates, and failure modes across various levels of triaxiality. Integrating stress-triaxiality-informed damage calibration into ballistic FE modeling represents an innovative approach of this study. Following this, FE models were validated and used to optimize multiple panel designs. FE modeling plays a crucial role in identifying promising panel configurations based on performance metrics, including perforation resistance, areal density (product of density and thickness), and cost. Promising panel configurations were then manufactured and single-hit impact tested at the center of the panel using a gas gun projectile launcher with STANAG Level 3 projectiles. Their ballistic performance and areal density were compared with those of the reference panels. The data from this study, based on the single-hit impact test, will enhance the performance of ballistic panels for GCV applications.

Materials and experimental procedure

Materials

The ballistic panels of the proposed protective system consist of three different materials. The first component is made from ballistic steel materials, supplied as plates with thicknesses ranging from 4 to 9 mm. Two grades of ballistic steel were used: BS550 and BS600, with nominal Brinell hardness values of 550 and 600 HBW, respectively. In addition to the steel layers, a UHMWPE composite was included. The commercially available UHMWPE composite was supplied as pre-consolidated plies, each consisting of four unidirectional layers with fibers arranged in a cross-ply [0⁰/90⁰]₂ orientation. The UHMWPE fibers are embedded in a polyurethane-based matrix with approximately 17% matrix by weight. The B₄C ceramic materials were obtained as hexagonal tiles produced by sintering.

Mechanical characterization

The effect of high strain rates and temperatures on the mechanical behavior of UHMWPE composites was investigated and reported in our previous study by Aziz et al. (2025). This continuation of the research details the specimen preparation and testing procedures used for the mechanical characterization of ballistic steel materials in the following subsections. The steels were tested at strain rates ranging from 10^-3 s^-1 to 1000 s^-1 and temperatures from 23°C to 300°C.

Specimen preparation

The mechanical properties of BS550 and BS600 needed for FE simulation were obtained from the uniaxial tensile test of a smooth dog-bone specimen, as shown in Figure 1(a). To study failure modes at different triaxiality levels, specimen geometries were specifically designed for each triaxiality level, as illustrated in Figure 1. For specimen preparation, the as-received ballistic steel plates were ground to a thickness of 1.8 mm, and the specimens were cut using wire Electrical Discharge Machining (EDM). Before testing, the specimens were coated with a thin layer of white paint, followed by black speckles, to create a speckle pattern for 3D Digital Image Correlation (DIC) displacement tracking. Three replicate specimens were tested for each condition.

Figure 1.

Different geometries of the specimens are used to cover various stress triaxialities in tensile testing. All dimensions are in mm.

Quasi-static test

Uniaxial tensile tests were conducted using an Instron 6800 UTM with a maximum capacity of 100 kN, equipped with a temperature chamber, as shown in Figure 2(a). Smooth-edge dog-bone specimens were used for testing and securely clamped with the UTM grip, as depicted in the detailed view in Figure 2(a). To enable full-field strain measurement within the effective gauge length of the specimens using the 3D-DIC technique, two 5.0-megapixel cameras from Correlated Solutions were calibrated and synchronized with the UTM. This non-contact system operates at two frames per second (fps) and can capture high strain levels, including near-failure conditions. Post-processing of the strain data was performed using Correlated Solution VIC-3D software. The specimens were loaded at a crosshead speed of 1 mm/min and tested until fracture. For tests conducted at temperatures of 25°C and 300°C, a 10-min thermal equilibration period was applied before testing to ensure the specimens reached the desired temperature. For these tests, the Lagrangian longitudinal strain, $ε_{x}$ , was measured from the DIC data. The material’s true stress, $σ_{t}$ , was calculated using the axial force, $F$ , and the specimen’s thickness, $t$ , and width, $w$ , and Lagrangian transversal strain, $ε_{y}$ , which was also measured using the DIC technique, as given in equation (1).

σ_{t} = \frac{F}{t w (1 - ε_{y})}

(1)

Figure 2.

(a) UTM with temperature chamber and DIC, (b) SHPB with DIC.

High strain rate tests

Dynamic uniaxial high-strain-rate tensile tests were conducted at room temperature using the SHPB Kariem et al. (2019), Ghisi et al. (2022). In these tests, the tensile specimens were placed between two cylindrical metallic bars (the incident and transmitted bars) made of Maraging Steel C350, as shown in Figure 2(b). During the test, a 250 mm striker bar is driven by a gas gun, strikes the incident bar, and generates an elastic pulse that travels along the incident bar, $ε_{i} (t)$ until it reaches the opposite end. The pulse is partially reflected, $ε_{r} (t)$ , and partially transmitted, $ε_{t} (t)$ . Strain histories were recorded from the transmitted and incident bars using strain gauges positioned at the center of each bar. The output voltage was amplified using a 2310B Signal Conditioning Amplifier (Micro-Measurements, U.S.) and recorded with a PicoScope 5442D oscilloscope (Pico Technology, U.K.) at a sampling rate of one million samples/sec. Additionally, the SHPB tensile test setup was monitored with an ultra-high-speed camera, Kirana 7M (Specialised Imaging, U.K.), which captured synchronized images during the tests. These images were used to extract the strain field deformation on the specimen surface using the DIC technique. The details of the test formulation and technique used are provided in our previous work Ghisi et al. (2022). The equations for determining stress $(σ (t))$ , strain ( $ε_{s} (t)$ ), and strain rate ( ${\dot{ε}}_{s} (t)$ ) of the tested specimen are given by Alves (2024), Ramos et al. (2022).

ε_{i} (t) = ε_{t} (t) - ε_{r} (t)

(2)

σ (t) = E_{b} \frac{A_{B}}{A_{s}} ε_{t} (t)

(3)

ε_{s} (t) = \frac{2 C_{0}}{l_{s}} \int_{0}^{t} ε_{r} (t) d t

(4)

{\dot{ε}}_{s} (t) = \frac{2 C_{0}}{l_{s}} ε_{r} (t)

(5)

where

E_{b}

is the Young’s modulus of the bar and

l_{s}

is the length of the striker bar (250 mm).

Failure mode at various stress triaxiality levels

The failure mode and failure strain were measured at various stress triaxiality levels, $η$ , using different specimen geometries, as shown in Figure 1. Different specimens will exhibit varying fracture properties and macroscopic failure modes. Distinct fracture mechanisms, such as shear-dominated, stretch-dominated, and mixed modes, can be analyzed using a triaxiality-based failure criterion Dresch et al. (2024). This assessment aimed to evaluate the material’s failure behavior under void coalescence ( $η > 0.33$ ) and shear-dominated ( $η < 0$ ) failure modes. These tests were conducted using UTM and SHPB, as detailed in previous sections. To characterize the deformation behavior, principal strains were measured at the center of the specimens. The methodology applied is presented in reference Sartor et al. (2022).

Development of ballistic panel configurations for STANAG level 3

Finite element modeling details

The stress triaxiality testing specimens were modeled using the Abaqus Explicit commercial FE software, with an element size of 0.2 mm in the fracture area. The failure strain from experimental triaxiality tests was applied in the ductile damage models for BS550 and BS600.

The impact phenomena were modeled using explicit FE in Abaqus, with the ballistic panels and projectiles simulated. For experimental testing, 7.62 mm x 51 AP-WC (M993) and, instead of API-B32, the equivalent projectile, 7.62 mm x 51 NATO AP (M61), were used due to limited availability and supply of projectiles. For the rest of the manuscript, M61 was referenced instead of API-B32, and M993 was referenced instead of AP-WC. Moreover, the Vickers hardness of polished projectile cores for M61 and M993 was measured at a 0.5 kgf load using the INNOVATEST NEMESIS 5100G2 instrument. Hardness measurements were taken at five points, and the average values were 749.21 ± 18.80 HV and 1413.72 ± 23.10 HV for the M61 and M993 projectile cores, respectively. The projectile details are given in Table 1, which were used to model the M61 and M993 projectiles.

Table 1.

Projectile types and dimensions.

Projectile	Projectile dimensions
Projectile	Length (mm)	Diameter (mm)	Weight (g)	Core material
7.62 x 54 R B32 API	37.8	7.8	10	Hardened steel
7.62 mm x 51 mm AP (M61)	32.95	7.82	9.75	Hardened steel
7.62 mm x 51 AP-WC (M993)	29	7.84	8.3	Tungsten carbide (WC)

The reference ballistic steel panel configuration (RBP1) - (BS600)_{5.5 mm}/(air gap (AG))_{15 mm}/(BS550)_{8.5 mm}), as shown in Figure 3(a) were modeled using hexahedral elements. The M61 and M993 projectiles were modeled using tetrahedral elements as shown in Figure 3(b) and (c), respectively. The Materials (hardened steel, WC, brass, and lead) in the projectiles were simulated using the Johnson-Cook plasticity and damage model, with their parameters taken from the literature ArcelorMittal (2024), Jones (2011). The BS550 and BS600 used the Johnson-Cook plasticity model and the Abaqus ductile damage model. The properties of BS550 and BS600 were obtained through experimental tests. Based on mesh sensitivity analysis and element size compatibility considerations, the element sizes for the metal plate, ceramic tile, and UHMWPE layer in the impact region were set to 0.25 mm, 0.25 mm, and 0.5 mm, respectively. The element size for the projectiles was set to 0.5 mm. The erosion criterion for each material was defined to match the corresponding mesh size. The outer edges of the ballistic panel’s back face were fixed, and an initial velocity was applied to the projectile, as specified by STANAG 4569. Automatic general contact was established between all entities. The UHMWPE materials were modeled using a proprietary user-defined subroutine, which sourced material properties from Aziz et al. (2025). The subroutine explicitly accounts for several critical deformation and failure mechanisms in UHMWPE laminates. It accounts for the strain-rate dependence of both tensile strength and axial (longitudinal) moduli. Fiber tension, compression, and crush failure modes, matrix shear failure mode, and matrix delamination failure mode are included in the model. Delamination of the laminate was modeled using a Cohesive Zone Model (CZM) with a bilinear traction-separation law to represent interlaminar failure between sub-laminates. The primary analyses in the FE model included evaluating perforation capability, examining the role of materials in the ballistic panel, and determining the ballistic limit.

Figure 3.

Modeling of the (a) reference ballistic steel panel 1 (RBP1); and projectiles: (b) M61 and (c) M993.

In FE simulations, the B₄C ceramic tiles were modeled using the JH-2 model, a widely used phenomenological model for brittle materials such as ceramics under high-strain-rate, high-pressure ballistic loading. The JH-2 model captures pressure-dependent strength, strain-rate effects, and progressive damage accumulation characteristic of brittle ceramics under ballistic impact. It assumes that the damage variable increases gradually with plastic deformation. A failure strain was defined to trigger element erosion.

Manufacturing of ballistic panel configurations based on FE

Initially, a reference ballistic steel panel 1 (RBP1) - ((BS600)_{5.5 mm})/(AG)_{15 mm/}(BS550)_{8.5 mm}) was used to evaluate the proposed configuration. The reference panel consists of two different high-hardness BS plates. The back face plate is 8.5 mm thick of BS550, while the front face plate is 5.5 mm of BS600. The 160 mm x 160 mm panels were assembled with an AG of 15 mm between them, as shown in Figure 4(a). The total areal density of the RBP1 is 132.6 kg/m². RBP 1 is intended to protect the M61 projectile at a velocity of 854 m/s. All panels manufactured to test against M61 were designed based on FE simulation and RBP1. From the literature [ArcelorMSittal (2024)], it was reported that the panel (BS600)_{9 mm}/(AG)_{15 mm}/(BS380)_{13.5 mm} with the areal density of 177 kg/m² could withstand the M993 projectile. Therefore, this panel (BS600)_{9 mm}/(AG)_{15 mm}/(BS380)_{13.5 mm} is considered the reference ballistic steel panel 2 (RBP2) against the M993 projectile. All panels manufactured to test against M993 were designed based on FE simulation and RBP2.

Figure 4.

Schematic representation of the ballistic panel configurations: (a) RBP1; (b) BP1; (c) BP2 and (d) BP3.

Following FE optimization, three panel configurations were developed to improve ballistic performance while reducing overall weight. The optimization process involved refining the material thickness and hybridizing the ballistic steel plates by strategically adding high-performance components, including UHMWPE and B₄C ceramic layers. The UHMWPE composite laminates were produced by compression molding on a Dishaa hydraulic hot press with a 300-ton capacity. First, the UHMWPE sheets were cut to the required dimensions and stacked in a [0⁰/90⁰]_2n sequence. The stacked sheets were placed between two flat steel platens and a Diatex fluoropolymer release film to prevent sticking to the steel mold. The material was consolidated at a temperature of 125°C and a pressure of 30 MPa, following the manufacturer’s recommended cycle Aziz et al. (2025). Throughout the cooling process, the pressure was maintained until the temperature fell below 50°C.

To improve ballistic resistance, B₄C ceramic was placed at the strike face to shatter and absorb the projectile’s kinetic energy upon impact. For ceramic-based setups, mosaic ceramic ballistic plates were employed, made of tiles arranged side by side. Figure 4(b)–(d) illustrates the optimized configurations of panels BP1, BP2, and BP3, with total areal densities of 103.6, 97.8, and 112.0 kg/m², respectively. The thickness of these panels is not disclosed due to confidentiality in our design. However, the impact performance of the developed panels, based on their areal density, is compared with that of the reference panels (RBP1 and RBP2). Areal density provides a consistent basis for comparing the performance of panel configurations.

Experimental testing of ballistic panel configurations

Experimental impact tests were conducted using a gas-gun projectile launcher, as shown in Figure 5(a). Compressed helium gas was employed to accelerate M61 and M993 projectiles to velocities of up to 1000 m/s. Initially, the projectile is placed inside a sabot, which is then propelled through the barrel. Once the sabot exits the barrel, it separates from the projectile, allowing only the projectile to hit the panel. The panels were fixed against a flat anvil with a square opening measuring 120 mm by 120 mm, enabling open back-face deformation or complete perforation by the projectile, as shown in Figure 5(b). The system includes a laser barrier to measure impact velocity and high-speed cameras to capture both the front and rear faces of the panel. The impact velocity and back-face displacement are also measured using high-speed images analyzed with the Photron Fastcam Analysis (PFA) tracking software and the Correlated Solutions VIC-3D DIC system.

Figure 5.

Gas-gun projectile launcher.

Results and discussion

BS550 and BS600 material characterization

Figure 6(a)–(d) shows the true stress–strain tensile curves of BS550 and BS600 at different temperatures and strain rates. Overall, both materials exhibit consistent behavior across repeated tests, confirming the reliability of the experimental results. Both materials exhibit elastic–plastic behavior and high strength. From Figure 6(a)–(b), it is clear that BS600 consistently outperforms BS550 in peak strength at room temperature, with stress levels exceeding 2200 MPa—much higher than typical steel alloys. Additionally, Figure 6(a)–(b) indicate that at high temperature (300°C), there is slight softening, reducing both yield and ultimate strength, especially in BS600, with about a 15% decrease. Overall, the temperature testing results show that the tensile samples become more ductile, with a significant increase in strain to failure. This aligns with studies on the dynamic softening of steels at elevated temperatures Cadoni et al. (2018).

Figure 6.

Reference materials mechanical characterization: temperature effect on

In Figure 6(c) and (d), the influence of strain rate becomes more apparent. At high strain rates (1753 s^-1 for BS550 and 1352 s^-1 for BS600), both materials show increased peak strength by approximately 200–300 MPa compared to their quasi-static conditions (0.003 s^-1). Although BS600 exhibits greater strength, BS550 demonstrates better ductility and specific energy absorption. The results indicate that BS600 may be more suitable for the strike face, as it can fragment and erode the projectile on impact due to its higher strength and nominal hardness. Conversely, BS550 might be better suited for the back face of the panel, because it can absorb residual impact energy and withstand remaining fragments after projectile perforation. This observation aligns with the findings of Singh et al. (2021), who noted that higher-strength steels exhibit increased resistance to penetration but may lose ductility at higher loading rates. Similarly, Ren et al. (2018) demonstrated that martensitic low-alloy steels maintain high strength over a wide range of strain rates, from quasi-static to 10⁶ s^-1, further supporting the applicability of BS600 in high-strain-rate environments. By obtaining the fracture strain and the corresponding average stress triaxiality for each tested specimen (Figure 1), the representative fracture locus was constructed and shown in Figure 6(e)–(f). The average stress triaxiality was computed from finite element simulations in which only the plasticity model was defined, with no damage model included. Elements in the central region of the specimen were selected, and their average stress triaxiality over time was computed. The stress triaxiality at the plateau stage of the stress triaxiality-time curve was then averaged. These triaxiality values correspond to the initiation of damage.

It is evident that fracture strain varies significantly with stress triaxiality for both BS550 and BS600 steels. Conventional tensile tests conducted on notched round specimens (Figure 1(a)) revealed that fracture consistently initiated at the center, where both the equivalent strain and stress triaxiality were highest, typically around $η \approx 0.4$ Bao and Wierzbicki (2004). These tests are representative of ductile fracture behavior under high triaxiality conditions. In contrast, tests on smooth dog-bone specimens (Figure 1(b)) resulted in fracture initiation at void coalescence, with a corresponding triaxiality of η ≈ 0.33. These specimens offer insight into the response of hardened steel in more moderate triaxiality regimes. Finally, in the pure shear configuration (Figure 1(c)), fracture occurred under nearly zero hydrostatic pressure, corresponding to $η < 0$ , where the fracture strain was significantly higher due to the dominance of shear-driven deformation mechanisms.

From the resulting fracture strain data, a logarithmic trendline was observed across all specimens, also highlighted in dashed lines, confirming the inverse relationship between triaxiality and ductility, a pattern widely reported in literature Rickhey and Hong (2022), Bao and Wierzbicki (2004), Erice et al. (2012). Under quasi-static loading, BS550 and BS600 exhibited fracture strains ranging from approximately 0.08 mm/mm to 0.12 mm/mm in high-triaxiality tensile tests, and significantly higher values in shear-dominated tests. This reinforces the notion that ductile fracture is highly stress-state-dependent, particularly under varying failure mechanisms, such as void coalescence and shear localization Erice et al. (2012).

In addition to triaxiality effects, the influence of temperature and strain rate was also evaluated using both smooth tensile and shear specimens. As illustrated in Figure 6(a)–(d), fracture behavior shifts towards a more brittle response at high loading rates. Conversely, elevated temperatures enhanced ductility, delaying failure and increasing strain to fracture, especially notable in BS550, which inherently has a higher elongation capacity. These findings highlight the importance of damage evaluation for numerical modeling, as ignoring the effects of triaxiality, particularly under dynamic conditions, can significantly compromise reliability. High triaxiality levels hinder plastic deformation and accelerate void growth, whereas low or negative triaxiality conditions enable extended shear deformation and higher fracture strain. Empirical models such as the Johnson–Cook failure criterion incorporate triaxiality in a simplified manner, but their accuracy often diminishes outside calibrated conditions, particularly in intermediate or mixed-mode regions (0 < η < 0.33), despite also accounting for temperature and strain rate effects Rickhey and Hong (2022). However, the transitional behavior between dominant failure modes is still not well understood and may involve complex interactions between local microstructure, strain rate sensitivity, and constraint effects.

The Johnson-Cook plasticity model was used to simulate the deformation behavior of the hardened materials. Table 2 summarizes the material parameters obtained from experimental observations. In our simulations, an inelastic heat fraction (with a default value of 0.9) was specified for metallic materials, which convert plastic work into heat, resulting in a local temperature rise. Since the J-C plasticity model explicitly includes a thermal softening term, thermal softening occurs whenever the temperature rises. Therefore, thermal softening during impact is captured within the standard J-C plasticity formulation. We acknowledge that we account for thermal softening here by extrapolating quasi-static high-temperature data to dynamic impact scenarios. However, adiabatic heating during impact is strain-rate dependent, which may lead to some discrepancies. Despite this limitation, the Johnson-Cook plasticity model has been widely and successfully used in ballistic impact simulations for many metallic materials.

Table 2.

Material parameters used for the FE modeling of BS550 and BS600.

Mechanical properties	BS550	BS600
Density (kg/m³)	7850	7850
Young’s modulus (GPa)	201	201
Poisson’s ratio	0.33	0.33
A (MPa)	1425.28	1455.00
B (MPa)	1059.67	1768.00
n	0.3	0.3
C	0.0117	0.0096
m	1.635	0.972

Abaqus ductile damage criterion model was used to simulate the damage behavior of the hardened materials, predicting the initiation of damage through nucleation, growth, and coalescence of voids. This damage model allows us to define it in tabular form, in which the equivalent plastic strain at the onset of damage is specified as a function of stress triaxiality (η), strain rate, and temperature. This approach doesn’t rely on the specific functional form of the Johnson–Cook damage equation, which is known to be sensitive to the calibration path and to exhibit limited Lode-angle dependence, particularly under mixed-mode fracture conditions (0 < η < 0.33). Alternative models, such as modified J-C, GISSMO, or advanced Lode-angle-dependent formulations, were considered during the model selection phase. However, the Abaqus ductile damage model was chosen for BS550 and BS600. It’s highly flexible for defining ductile damage using tabular data and can accommodate non-monotonic fracture loci commonly found in advanced alloys. Additionally, it eliminates global fitting errors by using direct interpolation and allows independent treatment of stress triaxiality, strain rate, and temperature effects.

Figure 7(a)–(c) show the load-displacement responses for the smooth, notched, and shear specimens made of BS550 and BS600, comparing experimental data with FE simulation results. Across all three specimen types, the numerical predictions closely match the experimental curves, accurately representing the initial stiffness, peak load, and post-peak softening behavior. This consistency indicates that the plasticity and ductile damage criteria parameters used in the constitutive model were well-calibrated and effectively captured the material’s response under different stress conditions. Specifically, the simulations successfully replicated the transition from elastic to plastic deformation, as well as the gradual reduction in load associated with damage accumulation and fracture. Additionally, the experimental and numerical results showed good agreement, with deviations remaining within 10% for all specimens. This consistency between the data validates the FE model’s ability to capture the effects of geometry-induced triaxial stress, supporting further development of ballistic panels as conducted in this work.

Figure 7.

Comparison of the load-displacement curves between the experimental results and FE simulations of BS550 and BS600 for the (a) smooth; (b) notched and (c) shear specimens.

Figure 8 compares the strain field deformation at crack initiation and failure regions, obtained from experimental DIC results and FE simulations of smooth, notched, and shear specimens made of BS550 material. These comparisons evaluate the calibrated FE model’s ability to accurately predict local deformation and failure mechanisms across different stress states. In the smooth dog-bone specimen (Figure 8(a)), the strain distribution remains fairly uniform until necking begins at the center of the specimen. Both DIC and FE results show this behavior, with fracture occurring in the region of highest von Mises strain, which is typical of ductile failure under uniaxial tension. The notched specimen (Figure 8(b)) displays significant strain localization at the notch, where geometric stress concentration causes early damage initiation. The FE model closely matches the experimental strain field, demonstrating its ability to accurately capture triaxiality-driven localization and crack initiation within constrained geometries.

Figure 8.

Comparison of equivalent von Mises strains (mm/mm) and fracture regions between experimental results and FE simulations of BS550 for (a) smooth; (b) notched and (c) shear specimens.

Additionally, the shear specimen (Figure 8(c)) exhibits highly localized failure along the inclined shear band, indicating a shear-dominated failure mechanism. The fracture occurs along this band, and FE accurately reproduces both the orientation and extent of strain localization observed experimentally. This supports the FE model’s ability to predict failure under low or negative triaxiality, where ductility increases and the damage process is driven primarily by shear failure with adiabatic stress rather than void coalescence. Overall, the strong agreement between experimental DIC results and FE predictions across all specimen types highlights the importance of accounting for geometric variations to simulate different stress triaxiality states. This validation also demonstrates the reliability of the calibrated material model in capturing both global and local deformation behaviors under complex loading conditions, which is essential for precise ballistic panel design and impact prediction.

FE ballistic limit validation with experiments

Figure 9(a) presents the simulation of RBP1 impacted by the M993 projectile at 900 m/s. The outcomes demonstrate high consistency between the two methods. The FE model successfully identified localized brittle failure on the panel’s front face (BS600), followed by ductile deformation on the rear face (BS550). The projectile’s erosion pattern also closely aligns with experimental results. In Figure 9(b), the ballistic limit chart compares the reference and numerical predictions of RBP1 against the M993 projectile. The simulation estimated a ballistic limit of 570 m/s, while experimental results showed 580 m/s, resulting in a 1.2% discrepancy.

Figure 9.

FE model validation and verification of RBP1 against M993: (a) a series of plots showing the impact process at various time intervals and (b) comparison of FE-predicted ballistic limit and residual velocities with experimental results.

Optimizing ballistic panel configurations with FE simulations

The FE model was used to explore promising panel configurations and determine material thicknesses that meet STANAG 4569 Level 3 specifications. More than 40 configurations were simulated, leading to the selection of a few promising ballistic panel designs for manufacturing and testing against M61 and M993 projectiles. All configurations were screened using FE simulation results, based on perforation resistance, areal density, thickness, and cost. However, the primary goal was to minimize panel areal density while attaining the maximum impact performance. The analysis examined typical designed configurations with two steel plates, UHMWPE composites between metallic layers or at the back of the panels, ceramic layers on the strike face, and combinations of these materials, as shown in Figures 10 and 11.

Figure 10.

Typical ballistic panel designs tested with the M61 projectile during development and optimization include: (a) two steel plates; (b) UHMWPE sandwiched between steel plates; (c) UHMWPE at the back of the panel (spall liner) and (d) a ceramic layer on the strike face of the panel.

Figure 11.

Typical ballistic panel designs impacted by the M993 projectile during design and optimization:

Figures 10(a) and 11(a) show that the FE simulation results indicate that the RBP1 panel with an areal density of 132.6 kg/m² withstood the M61 projectile but was perforated by the M993 projectile. This confirms that more protection is required to withstand the M993 projectile than the M61 projectile. Similar findings were reported in other studies Klement et al. (2008). From Figure 10(b) and (c), it is observed that at the same areal density of 103.6 kg/m², the configuration with UHMWPE placed between the steel plates was perforated, whereas the configuration with UHMWPE at the back face withstood the impact. This indicates that UHMWPE is more effective as a back-face spall liner. It is clear from Figure 10(c) and (d) that to withstand the M61 projectile, either a thick UHMWPE back-face spall liner or a B₄C strike face is required. However, to withstand the M993 projectile, a combination of a B₄C strike face and a UHMWPE back-face spall liner is required, as shown in Figure 11(b).

Ballistic panels testing

The ballistic panel configurations, summarized in Table 3, were selected after completing the FE model design phase. These panels were experimentally tested against both M61 and M993 projectiles to meet the STANAG Level 3 requirement. Their performance and areal density were compared with those of RBP1 and RBP2, respectively. Initially, RBP1 with an areal density of 132.6 kg/m², was subjected to a single-hit impact at the center of the panel using both M61 and M993 projectiles. RBP1 withstood the M61 projectile but was perforated by the M993 projectile. This demonstrates that passing the M61 does not ensure survival against the M993. However, if the panel can withstand the M993 projectile, it will also withstand the M61 projectile.

Table 3.

Impact performance of ballistic panels against M61 and M993 projectiles.

Panel ID	Multi-layered configuration	Areal density (kg/m2)	Impact velocity (m/s)	Result
STANAG level 3 - M61 projectile with hardened steel core
RBP1 (steel core)	BS600/BS550	132.6	850.0	Withstand
BP1	BS600/BS550/UHMWPE	103.6 (22% lighter)	867.0	Withstand
BP2	Ceramic/BS600/BS550	97.8 (26% lighter)	900.2	Withstand
BP3	Ceramic/BS600/BS550/UHMWPE	112.0 (9% lighter)	Not required to be tested as it withstands M993	Withstand
STANAG level 3 – M993 projectile with WC core
RBP2 (WC core)	BS600/BS380 (Arcelormittal (2024))	177.0	930.0	Withstand
RBP1	BS600/BS550	132.6 (25% lighter)	925.0	Perforation
BP2	Ceramic/BS600/BS550	97.8 (45% lighter)	Not tested	Perforation (based on FE)
BP3	Ceramic/BS600/BS550/UHMWPE	112.0 (37% lighter)	933.3	Withstand

The results of three configurations, (i) BP1, (ii) BP2, and (iii) BP3, against the M61 projectile are given in Table 3 and discussed here. Notably, the BP2 configuration features a B₄C ceramic layer on the strike face, enabling thinner steel plates and reducing areal density compared with RBP1 and other panel configurations. All three panels withstood the M61 projectile, although their areal densities are 22%, 26%, and 9% lower than that of RBP1, indicating their superior specific energy absorption compared to RBP1. Based on FE predictions (Figure 10(c)) and further experimental validation (Figure 12(b)), the BP2 panel was the best choice against the M61 projectile in terms of specific energy absorption (i.e., good impact performance at lower areal density) compared with the other panels.

Figure 12.

Post-impact of the ballistic panels: (a) Comparison of FE predicted BFS and experimentally measured BFS using DIC; (b) BP1 against M61 projectile; (c) BP2 against M61 projectile and (d) BP3 against M993 projectile.

The results for two panel configurations, (i) BP2 and (ii) BP3, against the M993 projectile are discussed here. The areal densities of these panels are 45% and 37% lower than that of RBP2, respectively. Based on FE predictions, the BP2 panel could not withstand the M993 projectile and exhibited lower impact resistance than the RBP2 panel. Therefore, the BP2 panel was not experimentally tested against the M993 projectile. Based on FE predictions (Figure 11(b)) and further experimental validation (Figure 12(c)), the BP3 panel withstood the M993 projectile. Compared with the BP2 panel, the BP3 panel exhibited superior impact resistance, mainly attributed to the UHMWPE spall liner. As the BP3 panel withstood the M993 projectile, it is not necessary to test this panel against the M61 projectile.

Failure mechanisms of ballistic panel configurations

Figure 12(a) shows the FE-predicted and experimental back-face signature (BFS) for BP1 panel configuration against M61 projectile. The FE-predicted BFS of 24 mm was validated by an experimental BFS of 22 mm measured using DIC. To the best of the authors’ knowledge, STANAG Level 3 does not specify a BFS limit, and acceptable performance is defined as avoiding perforation. Strike-face damage in the tested panels of BP1, BP2, and BP3 is shown in Figure 12(b), 12(c) and 12(d), respectively.

Figure 12(b) shows the post-impact damage of the BP1 panel configuration with a BS600 strike face subjected to the M61 projectile. Initially, the projectile contacts the BS600 strike face, where a localized fracture characterized by ductile hole formation occurs in the steel plate. A similar failure mechanism was reported in the literature Ranaweera et al. (2023). The remaining projectile continues to interact with the UHMWPE layers, causing fiber shearing and localized perforations, which lead to failure in the initial portion of the UHMWPE layers Joshi et al. (2024). Following this, primary yarn stretching occurs within the remaining layers at the projectile impact zone. This deformation propagates in the out-of-plane direction, forming a cone in the rear layers. Secondary yarn deformation occurs in the surrounding regions, contributing to additional load distribution. These deformation mechanisms enable the UHMWPE material to dissipate residual kinetic energy, resulting in the observed back-face deformation, as shown in Figure 12(a).

In Figure 12(c), during the initial stage of projectile impact, the ceramic front surface experienced compressive loading, which induced localized radial and circumferential brittle fracture on the rear of the ceramic while eroding the projectile nose. The projectile jacket deformed during interaction with the ceramic, exposing the underlying steel core. In addition, the confinement of ceramic fragments increased projectile erosion. Subsequent abrasion mechanisms blunted and shortened the steel core, reducing projectile mass. These combined deformation and erosion processes increase the penetration resistance of the ceramic strike face panel configuration compared with that of the ballistic steel strike face panel configuration. The targeted tile in direct contact with the projectile was completely fractured. However, the damage remained confined, with only slight cracking in adjacent tiles. This localized damage pattern highlights the ceramic mosaic’s ability to preserve the integrity of surrounding tiles, thereby improving multi-hit performance.

For configurations incorporating ceramics such as BP3 in Figure 11(b) and 12(d), the impact resulted in both ductile hole formation and a bulge at the front face of the BS600 steel plate. A similar UHMWPE failure mechanism was observed, as previously described for Figure 12(b). The UHMWPE rear layers act as an effective spall liner, protecting the structure from perforation. Based on this single-hit impact test, the BP3 panel has been identified as the optimal configuration, offering superior ballistic performance and lower areal density compared to the other configurations. Therefore, multi-hit impact tests in accordance with the STANAG Level 3 standard are recommended to evaluate the ballistic performance of this configuration for GCV applications.

Conclusions

In this study, multi-layered ceramic/metal/UHMWPE composite ballistic panels were designed and developed to meet STANAG 4569 Level 3 protection standards. First, mechanical characterization was performed on various BS550 and BS600 specimen geometries to examine failure modes at different levels of stress triaxiality. The tests were conducted at strain rates from 10^-3 s^-1 to 1000 s^-1 and temperatures ranging from 23°C to 300°C, using 3D-DIC. Additionally, finite element simulations were performed to predict the mechanical properties and were validated against experimental data, showing excellent agreement. Furthermore, the reference ballistic panel 1 (RBP1) was experimentally tested to determine its ballistic limit. The ballistic limit of RBP1 was estimated using finite element analysis and experimentally validated, and both are in good agreement, with a deviation of less than 5%. Subsequently, three hybrid ballistic panel configurations were then fabricated based on reference panels and finite element simulations: (i) BP1 - (BS600/BS550/UHMWPE), (ii) BP2 - (B₄C/BS600/BS550), and (iii) BP3 - (B₄C/BS600/BS550/UHMWPE). These panels were tested against M61 and M993 projectiles at velocities up to 1000 m/s. Results from a single-impact test indicated that BP3 provides superior ballistic protection compared to all other panels and reduces areal density by 37% relative to the reference ballistic panel, RBP2, due to the combination of ceramics, metal, and UHMWPE. The BP3 panel configuration is recommended for further testing to ensure its multi-hit capability in accordance with the STANAG Level 3 standard for ground combat vehicle applications.

Footnotes

Acknowledgments

The work was supported by the Technology Innovation Institute (TII), Abu Dhabi, UAE.

ORCID iDs

Alia Ruzanna Aziz

Naresh Kakur

Rafael Savioli

Henrique Ramos

Funding

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

Declaration of conflicting interests

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

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