Performance of a Hybrid Composite of Kevlar, Aluminum and Cabuya Fiber Against Ballistic Threats—Numerical and Experimental Study

Abstract

The growing demand for lightweight and cost-effective vehicular armor systems has driven the development of hybrid multilayer architectures capable of improving ballistic resistance while reducing structural mass. This study evaluates the ballistic performance of a functionally graded aluminum–Kevlar–cabuya fiber composite system designed for vehicle door protection. A combined experimental–numerical framework was implemented, integrating ballistic testing according to NIJ 0108.01 and STANAG 4569 Level 1 standards with explicit dynamic finite element modeling based on the Johnson–Cook constitutive formulation for AA5083-H32. The multilayer configuration (25 mm aluminum/15 mm Kevlar 29/15 mm treated cabuya composite) successfully resisted 9 × 19 mm and 5.56 × 45 mm FMJ threats without complete perforation. Numerical simulations predicted a maximum back-face deformation of 52.75 mm under 9 mm impact, showing strong agreement with the experimental measurements (mean ± SD, n = 3). Post-impact microstructural analysis revealed a sequential energy dissipation mechanism governed by plastic deformation of the aluminum layer, Kevlar fibrillation and fragment retention, and controlled micro-cracking within the treated cabuya backing layer. With an areal density of 140.87 kg/m², the system achieved a 19% weight reduction compared with conventional steel-based solutions. These results demonstrate the structural-scale feasibility of integrating treated cabuya fiber composites as active energy redistribution layers in certified hybrid vehicular armor systems.

1. Introduction

Increasing insecurity in several regions of the world has intensified the demand for accessible and effective ballistic protection solutions, particularly in urban environments with high levels of gun violence. In Latin America, this situation has led to a growing demand for vehicle armouring systems in response to organised crime and armed attacks [1,2]. However, the implementation of these protective solutions faces two major challenges: the high cost of conventional armoring systems, which often exceed £17,000 per unit, and the lack of regulatory oversight in local markets, which may compromise the reliability and certified performance of available products [3,4].

Advanced materials such as titanium alloys, ceramics, high-performance polymers, and aluminum alloys have demonstrated excellent ballistic resistance in both military and civilian applications [5,6,7]. Nevertheless, their high production and processing costs significantly limit large-scale adoption, particularly in regions with restricted economic resources. Consequently, there is a growing need to develop cost-effective armor systems capable of maintaining certified ballistic performance while reducing both weight and manufacturing costs. One promising approach involves hybrid material configurations that combine natural fibers with high-performance synthetic reinforcements, such as Kevlar and aluminum alloys, to optimise the protection-to-weight and cost ratio [8]. Such hybrid solutions may therefore provide a viable alternative for vehicular armor in high-risk environments.

The development of hybrid armor systems has attracted considerable attention due to their ability to exploit the complementary mechanical properties of different materials, thereby enhancing ballistic efficiency while reducing overall structural mass. Recent studies have shown that combining Kevlar and UHMWPE fibers with aluminum alloys and natural fiber composites can provide superior protection compared with traditional monolithic steel armor systems [8,9]. Within this context, the incorporation of natural fibers such as cabuya (Furcraea andina) represents an emerging strategy for introducing lightweight and sustainable reinforcement layers into graded ballistic architectures [10,11]. Cabuya fiber is characterized by longitudinally aligned cellulose microfibrils embedded in a lignin–hemicellulose matrix, which promotes favourable specific mechanical properties and progressive fibrillation under tensile loading. Owing to its relatively low density (≈1.3–1.5 g·cm⁻³), cabuya contributes to mass reduction in multilayer systems while also facilitating energy dissipation mechanisms such as fiber pull-out, interfacial debonding, and controlled delamination [12]. Unlike synthetic high-performance fibers, cabuya can be cultivated with limited industrial processing requirements and is widely available in Andean regions, supporting its scalability as a reinforcement material. Although its absolute tensile strength is lower than that of aramid or UHMWPE fibers, its integration as a secondary or backing layer in hybrid armor configurations may enhance stress redistribution after projectile deceleration by the frontal metallic and high-tensile layers. This graded mechanical response enables the natural fiber composite to function not as a primary penetration barrier but rather as a structural energy-absorbing and deformation-controlling component within the multistage ballistic mitigation process. For example, carbon/Kevlar hybrid laminates have demonstrated improved impact tolerance and damage mitigation by combining the high stiffness of carbon fibers with the superior energy absorption capacity of ductile aramid fibers [13,14]. Similarly, the incorporation of natural fibers such as sisal and cabuya has emerged as a promising research direction in ballistic composites, with studies indicating their potential to reduce structural weight while offering sustainable and economically accessible reinforcement alternatives [15]. Hybrid armor structures integrating materials with complementary mechanical responses have demonstrated enhanced penetration resistance and improved energy dissipation mechanisms [16]. In particular, the stacking sequence and layer arrangement play a critical role in determining ballistic performance. Configurations combining rigid frontal layers (ceramics or metals) with ductile rear reinforcements (Kevlar or UHMWPE) have proven especially effective in mitigating penetration and reducing back-face deformation [17,18]. These systems have been validated through combined experimental and numerical approaches, confirming their effectiveness against diverse ballistic threats [19].

Experimental and numerical analyses play a fundamental role in the evaluation and optimisation of hybrid armor systems. The finite element method (FEM), coupled with advanced constitutive models such as the Johnson–Cook formulation, enables accurate simulation of material behaviour under high strain-rate ballistic impact conditions [20]. Numerical modeling allows the prediction of stress distribution, plastic deformation, and energy absorption within complex multilayer architectures, significantly reducing reliance on costly experimental testing [21]. However, challenges remain in accurately validating numerical models for multiphase composite systems, particularly those incorporating natural fibers. Recent research highlights the need to improve modeling strategies capable of capturing complex interfacial phenomena, including delamination, fiber–matrix interaction, and stress-wave transmission across heterogeneous layers [22,23].

Despite significant advances in aluminum-, aramid-, and ceramic–polymer hybrid armor systems, most existing configurations rely primarily on dual-material strategies focused either on ductile metal deformation or on high-tensile fiber rupture. Only limited studies have explored tri-material graded architectures that incorporate natural fibers as functional structural backing layers rather than passive fillers.

Furthermore, much of the existing research focuses mainly on penetration resistance without fully analysing interfacial stress redistribution and multi-stage energy absorption mechanisms across heterogeneous layers. The main scientific gap addressed in this study is the lack of experimentally validated numerical models capable of predicting sequential energy dissipation mechanisms in aluminum–aramid–cabuya fiber hybrid systems specifically designed for vehicular ballistic applications. Existing studies typically evaluate natural fiber composites independently or combine synthetic fibers with metals without quantifying the synergistic effect of graded layer sequencing.

Therefore, this research proposes a functionally graded trilayer hybrid architecture in which each material layer performs a distinct mechanical role: (1) metallic plastic deformation and projectile blunting in the aluminum layer, (2) high-tensile fiber energy absorption through fibrillation in the Kevlar reinforcement, and (3) stress redistribution and controlled delamination through the treated natural fiber composite backing.

The research gap addressed in this study lies in the underutilisation of natural fibers, such as cabuya, in certified hybrid armor systems for high-security vehicular applications. While synthetic and metallic materials have been widely investigated, natural fibers remain insufficiently explored in multilayer ballistic configurations despite their advantages in terms of sustainability, cost efficiency, and density reduction.

The objective of this study is to evaluate the ballistic response of a functionally graded aluminum–aramid–cabuya hybrid armor system designed for vehicular protection. It is hypothesised that incorporating a treated cabuya fiber composite as a backing layer enhances multi-stage energy absorption through stress redistribution and controlled delamination, without compromising certified ballistic resistance.

This hypothesis is evaluated through a combined experimental–numerical framework integrating standardized ballistic testing and finite element modeling to characterise impact behaviour and interfacial energy transfer. Unlike previous studies focused on flat composite coupons or textile-dominated soft armor systems, this work provides experimentally validated door-scale performance of a functionally graded aluminum–aramid–cabuya fiber architecture. The analysis is supported by explicit dynamic finite element modeling capable of predicting sequential multi-material energy dissipation mechanisms.

2. Materials and Methods

2.1. Armor System Design and Numerical Modeling

The scientific methodology followed the workflow illustrated in Figure 1. The design of the multilayer armor system began with the precise acquisition of the door geometry of a Peugeot 3008 vehicle. Using high-precision measurement tools (digital calipers and tape measures), the critical dimensions—height, width, and thickness—were recorded to generate a parameterized three-dimensional CAD model in Autodesk Inventor Professional 2025. Non-structural components were excluded from the model to simplify the numerical analysis [16].

The design concept is based on an external armor plate adapted to the internal structure of the vehicle door. The configuration integrates structural reinforcements to optimise impact resistance while prioritising a secure anti-vibration fastening system. The final geometry includes the vehicle door (modelled as ST37 steel), a multilayer armor configuration consisting of AA5083-H32 aluminum, Kevlar reinforcement, and a natural fiber composite backing layer, as well as the projectiles represented as hardened steel cores used for penetration simulations [17,18].

The natural reinforcement selected for this study consisted of Furcraea andina fibers obtained from the Intag region (Imbabura, Ecuador). The fibers were embedded in a thermosetting epoxy resin matrix (commercial grade, room-temperature curing system) to fabricate the natural fiber composite backing layer. The epoxy system was selected due to its high interfacial adhesion with lignocellulosic fibers, adequate tensile strength, and proven performance in structural composite applications. The composite laminate was manufactured using a manual lay-up process followed by controlled curing at ambient temperature, ensuring uniform fiber distribution and consistent thickness across the 15 mm backing layer.

The polymer matrix used to fabricate the cabuya fiber composite consisted of a commercial two-component structural epoxy adhesive (3M™ Panel Bonding Adhesive 08115, 3M Company, St. Paul, MN, USA). This adhesive system is based on a toughened epoxy resin combined with a modified amine curing agent and is supplied in a dual-cartridge configuration that ensures automatic mixing during application. The base component has a density of approximately 0.96 g·cm⁻³, while the hardener exhibits a density of approximately 1.20 g·cm⁻³. The recommended mixing ratio is 172:100 by weight (or 200:100 by volume). The adhesive provides a working time of approximately 90 min at 23 °C and reaches handling strength after about 90 min, with full curing occurring after approximately 24 h at room temperature. The system exhibits high adhesion to metallic and composite substrates, with reported overlap shear strengths of up to approximately 27 MPa on steel and about 21.7 MPa on aluminum substrates. This epoxy system was selected due to its high bonding strength, corrosion resistance, and suitability for structural bonding in automotive body panels.

As illustrated in Figure 1, the armor layers are arranged in a functionally graded configuration. The first layer, shown in lead colour, consists of AA5083-H32 aluminum, which acts as the primary impact interface by absorbing the initial kinetic energy through strain-rate-dependent plastic deformation and promoting projectile blunting. The second layer, shown in yellow, consists of Kevlar 29 fabric, whose function is to absorb residual energy through tensile membrane stretching, fiber fibrillation, and progressive delamination mechanisms. The third layer, shown in orange, is composed of a treated cabuya fiber composite, which provides additional structural rigidity while redistributing residual stresses through controlled micro-cracking and limited interfacial damage, thereby mitigating excessive back-face deformation.

The specific combination of aluminum, Kevlar, and cabuya fiber was selected following a graded impedance and complementary failure-mechanism strategy. Rather than relying on a single dominant energy absorption mechanism, the architecture promotes sequential energy dissipation, consisting of metallic plastic work in the frontal layer, tensile energy absorption in the aramid reinforcement, and structural stress redistribution in the natural fiber backing. This tri-material configuration differentiates the proposed system from conventional dual-layer metallic–aramid armours by introducing an additional energy redistribution stage at the structural scale while maintaining weight and cost efficiency. The hybrid configuration was exported in STEP format to ensure compatibility with the simulation environment.

The model was then imported into ANSYS Workbench 2025 R1 for explicit dynamic simulation. A structured MultiZone meshing strategy, combining quadrilateral and triangular elements, was implemented to ensure high mesh quality and numerical stability during the impact simulations. Local mesh refinement was applied in the projectile–armor interaction region using body-sizing control, resulting in a minimum element size of 1 mm in the impact zone in order to accurately capture the high stress gradients generated during ballistic contact. The final mesh consisted of 17,555 elements and 13,581 nodes, with an average Element Quality of 0.935, ensuring reliable numerical performance of the explicit dynamic analysis [19,20]. The material properties of each component (AA5083-H32 aluminum, Kevlar, cabuya fiber composite, and steel) were defined based on the MatWeb material database [24] and specialised technical literature.

2.2. Material Characterization and Properties

The key mechanical properties of the constituent materials used in the hybrid shielding system (Figure 2a) were obtained from the MatWeb technical database and specialised literature, ensuring reliable input parameters for the computational model.

Aluminum alloy AA5083-H32 was selected for the front layer due to its combination of moderate strength, high ductility, and its well-documented use in both military and civilian ballistic shielding applications [22].

In this study, Kevlar 29 composite was selected as the intermediate layer for energy absorption because of its proven ability to resist penetration and dissipate the kinetic energy of projectiles. A natural cabuya fiber composite was incorporated as a reinforcement and backing layer, providing additional structural support to the armor system while also contributing to the sustainability of the material. Figure 2a–c illustrates the armor system model, where the Kevlar 29 layer and the natural fiber composite layer are arranged in a strategic sequence designed to optimise energy absorption and enhance the overall structural performance of the armor.

The door substrate was modelled as ST37 steel, equivalent to S235JR [17] a material commonly used in structural applications due to its mechanical strength and widespread availability. The projectile core (Figure 2d) was defined as high-carbon steel, selected to represent a realistic penetration threat [18]. This material definition allows accurate simulation of the interaction between the projectile and the multilayer armor system, providing a realistic representation of the ballistic behaviour of the proposed configuration [25].

The material properties used in the finite element analysis (FEA) are summarised in

Table 1. Mechanical properties of the materials used in the simulation.

Material	Density/(kg/m3)	Modulus of Elasticity/(GPa)	Coefficient of Poisson	Tensile Yield Strength/(MPa)	Ultimate Tensile Limit/(MPa)
Door—Steel ST37	7850	210	0.33	235	510
Armos—Aluminum 5083-H32	2660	70	0.33	250	320
Armor—Kevlar	1440	70 *	0.34	-	3600 *
Armor—Cabuya Fiber [23]	1300	0.0075 *	-	-s	305 *
Projectiles—(AISI 4340)	7850	192	0.29	470	745

Note: For composite materials (Kevlar, Cabuya Fiber), the anisotropic properties were simplified to an equivalent isotropic model for the initial analysis. Values marked with * correspond to the fiber direction.

. These parameters are essential for defining the material constitutive models implemented in ANSYS and for accurately predicting the structural response under high-velocity impact conditions.

The stacking sequence was designed following a graded impedance principle, in which materials are arranged from high plastic deformability (AA5083-H32 aluminum) to high tensile strength (Kevlar 29), and finally to a natural fiber composite layer that promotes energy dissipation through micro-cracking and delamination. This graded impedance configuration aims to minimise stress-wave reflections at the interfaces, thereby improving the overall ballistic efficiency of the multilayer system.

The numerical simulation model was developed using the finite element method (FEM) implemented in ANSYS Workbench. The Johnson–Cook constitutive model was applied to AA5083-H32 aluminum to simulate plastic deformation under high strain-rate conditions. The boundary conditions were defined to represent a rigid mounting system, ensuring that the projectile’s kinetic energy was primarily absorbed by the armor layers. In addition, a mesh convergence analysis was conducted by adjusting the mesh density until consistent results were obtained, with variations in maximum stress values remaining below 5%.

The configuration of the hybrid armor layers used in this study to protect the vehicle door is illustrated in Figure 3. On the left side of the figure, the physical armor sample is shown, where the different layers are clearly identified and arranged according to the experimental design.

• Aluminum layer 5083-H32 (marked in yellow): This layer acts as the first line of defence, providing the initial resistance to projectile impact. Aluminum was selected due to its favourable combination of strength, ductility, and energy absorption capability through strain-rate-dependent plastic deformation.
• Kevlar layer (marked in red): Positioned beneath the aluminum layer, the Kevlar reinforcement acts as an energy-absorbing layer that dissipates the residual kinetic energy of the projectile through fiber stretching, fibrillation, and progressive delamination. Kevlar is widely recognised for its exceptional impact resistance and high energy absorption capacity.
• Cabuya fiber layer (marked in orange): Located in the innermost region of the armor system, this layer consists of a cabuya fiber composite that contributes to residual stress redistribution and provides additional structural reinforcement. Furthermore, the incorporation of cabuya fibers enhances the sustainability of the system while contributing to overall weight reduction, which is particularly important for vehicular armor applications.

On the right side of the figure, the vehicle door made of ST37 steel is shown, forming the structural base of the armor system. This image provides a visual representation of how the armor layers were arranged and integrated into the internal structure of the door to enhance protection against ballistic impacts.

2.2.1. Constitutive Modeling of Composite Layers

Kevlar 29 fabric and treated cabuya fiber composite layers exhibit inherently anisotropic mechanical behaviour due to fiber orientation, weave architecture, and strain-rate sensitivity under ballistic loading conditions. Aramid fibers are characterised by high tensile strength along the fiber direction and significantly lower transverse and shear stiffness, resulting in orthotropic material symmetry. Similarly, woven cabuya fiber composites exhibit direction-dependent stiffness and progressive failure modes governed by fiber pull-out, matrix cracking, and delamination mechanisms.

Under ballistic impact conditions (strain rates on the order of 10³–10⁴ s⁻¹), Kevlar fabrics also exhibit dynamic tensile strength amplification, as reported in experimental investigations on high-performance aramid laminates [26]. Therefore, an accurate representation of composite behaviour ideally requires orthotropic constitutive models coupled with strain-rate-dependent failure criteria.

In the present study, an equivalent homogenised formulation was initially implemented to ensure numerical stability and computational efficiency in the full vehicle-door explicit dynamic model. To justify this modeling simplification, orthotropic elastic properties of Kevlar 29 were compiled from the validated experimental literature and manufacturer technical data. These values were then used to derive the equivalent homogenised stiffness parameters implemented in the finite element model.

In large-scale explicit dynamic simulations involving complete structural assemblies such as vehicle doors, the implementation of fully orthotropic, strain-rate-dependent constitutive models for multiple composite layers significantly increases computational cost and may introduce numerical instability due to the large number of material parameters required. For this reason, a homogenized equivalent isotropic formulation was adopted in the present study as a first-order approximation of the composite behaviour. This approach allows the numerical model to efficiently capture the global structural response of the multilayer armor system while maintaining computational robustness during high-velocity impact simulations.

Kevlar 29 woven fabric was modelled assuming a balanced plain-weave architecture (0°/90°), which allows in-plane orthotropic symmetry (E₁ ≈ E₂). The longitudinal modulus represents stiffness along the fiber direction, while the transverse and shear moduli account for matrix-dominated deformation and inter-yarn shear mechanisms.

Although strain-rate-dependent orthotropic damage evolution models (e.g., Hashin or Puck criteria) were not explicitly implemented in the present numerical framework, the model was calibrated and validated against experimental ballistic tests conducted under NIJ 0108.01 and STANAG 4569 Level 1 conditions. The strong correlation between predicted and experimentally observed deformation patterns indicates that the homogenised approach provides reliable predictions of the global ballistic response, including penetration resistance and back-face deformation.

Nevertheless, it should be emphasised that the homogenised isotropic formulation introduces certain modeling limitations when representing fiber-reinforced composites under ballistic loading. In particular, the model does not explicitly resolve direction-dependent stiffness variations associated with fiber orientation, nor does it capture strain-rate-dependent orthotropic damage evolution within the composite plies, despite the orthotropic elastic and strength properties reported in

Table 2. Orthotropic elastic and strength properties of Kevlar 29.

Property	Symbol	Value	Units	Reference
Longitudinal Young’s Modulus	E1	83	GPa	[27,28]
Transverse Young’s Modulus	E2	7	GPa	[28]
Through-thickness Modulus	E3	5	GPa	[26]
In-plane Shear Modulus	G12	3.0	GPa	[29]
Transverse Shear Modulus	G23	2.5	GPa	[28]
Major Poisson Ratio	ν12	0.34	–	[27]
Longitudinal Tensile Strength	Xt	3600	MPa	[27]
Transverse Tensile Strength	Yt	80	MPa	[26]
Shear Strength	S12	100	MPa	[27]
Density	ρ	1440	kg/m3	[27]

As a consequence, local intra-ply failure mechanisms—including progressive fiber rupture, yarn pull-out, matrix cracking, and interlaminar delamination—cannot be represented with full physical fidelity. These phenomena are known to influence the detailed damage morphology of ballistic composites, especially at the mesoscale level.

2.2.2. Characterization of Natural Reinforcement: Cabuya Fiber

The natural reinforcement consisted of Furcraea andina (cabuya) fibers recognized for their considerable length and tensile capacity (Figure 4a). Fiber extraction was carried out through mechanical shredding of mature leaves using a motorized shredder (Figure 4b), followed by crushing and washing to eliminate pulp residues and surface impurities. The extracted fibers (average diameter ≈ 1.5 mm; Figure 4c) were subjected to an alkaline treatment using a 5 wt.% NaOH solution for 2 h at ambient temperature to enhance fiber–matrix interfacial adhesion by removing hemicellulose and increasing surface roughness [23]. After treatment, fibers were thoroughly rinsed to neutral pH and oven-dried at 60 °C for 24 h to remove residual moisture. The treated fibers were arranged in a 0°/90° flat weave configuration and impregnated with a thermosetting epoxy resin system using a manual lay-up process to fabricate the composite backing layer (nominal thickness: 15 mm). Curing was conducted at room temperature for 24 h, ensuring complete crosslinking of the matrix. The resulting composite exhibited an elongation at fracture of approximately 5% and a density of 1.3 g/cm³. Mechanical testing indicated comparable stress response in fiber and transverse directions, suggesting reduced anisotropic effects relative to untreated fibers and improved energy dissipation capability under ballistic loading [23].

This selection of natural fibers is supported by the literature: composites with analogous fibers (example: abaca) in longitudinal orientation have shown increases in tensile stress between 138–268% and significant improvements in elastic modulus, proving their effectiveness in high-performance applications [10,23]. Furthermore, natural fiber was chosen not only for its strength-to-weight ratio (reported tensile strength values of 305 MPa and modulus of elasticity of 7.5 MPa in the fiber direction) [23], but also as a sustainability strategy, incorporating a renewable, biodegradable, and low-cost material, thereby reducing the environmental footprint of the armoring [30].

Figure 4 illustrates both the untreated and treated fibers, showing the significant structural changes induced by the alkaline treatment, which resulted in an enhancement of the interfacial properties critical for energy absorption in the hybrid armor.

2.3. Experimental Ballistic Testing Setup

Experimental validation of the shielding system was conducted under standardised ballistic testing protocols to ensure the repeatability and reliability of the results. The tests were performed at an authorised ballistic testing facility (Figure 5), with the support and supervision of specialised military personnel. The armoured door assembly was rigidly anchored to a metallic support structure designed to minimise post-impact movement, ensuring that the projectile’s kinetic energy was primarily absorbed by the armor system rather than by displacement of the target. The target was positioned perpendicular to the projectile trajectory during all test configurations.

To simulate realistic ballistic threats, two types of firearms and ammunition—selected due to their prevalence in gun violence scenarios—were used. The first corresponds to a high-risk rifle threat, represented by an ARAD assault rifle (5.56 × 45 mm) (Figure 6a) using Full Metal Jacket (FMJ) ammunition. This configuration represents a high-energy threat with significant penetrating capability. The second corresponds to a medium-risk pistol threat, represented by a Glock pistol chambered in 9 × 19 mm Parabellum with FMJ ammunition, as shown in Figure 6b. This configuration represents a common ballistic threat typically encountered in close-range attacks. The experimental procedure was designed following the guidelines for the evaluation of ballistic materials established in NIJ Standard 0108.01 [31,32]. Additionally, the structural integrity criteria for the armor system were based on the requirements defined for light armoured vehicles in the NATO standardisation agreement STANAG 4569 (Level 1) [31,32].

Multiple shots were fired at intact armor panels from standardised test distances. For the rifle threat (5.56 × 45 mm FMJ), the shots were fired from a distance of 15 m. The initial projectile velocity (v₀) was obtained from the manufacturer’s technical specifications, corresponding to 920 m/s for the 5.56 × 45 mm FMJ cartridge and 360 m/s for the 9 × 19 mm FMJ cartridge [33].

Using the ballistic simulation module implemented in ANSYS Workbench, the projectile trajectory and aerodynamic deceleration were modelled to estimate the impact velocity. The model considered the projectile geometry and ballistic coefficient (BC), air density, and the distance between the firearm and the target. This analysis yielded estimated impact velocities of 850 ± 25 m/s at 15 m for the 5.56 × 45 mm calibre and 300 ± 15 m/s at 6 m for the 9 × 19 mm calibre.

The primary performance criterion was non-penetration of the armor system, in accordance with the protection requirements specified in NIJ Standard 0108.01 and STANAG 4569. After each impact event, a visual and tactile inspection of the panel was conducted to verify the structural integrity of the armor system. Additionally, the post-impact back-face deformation (BFD) on the inner side of the target was documented photographically to enable qualitative assessment of energy absorption and the associated failure mechanisms.

Statistical Analysis of Experimental Data

Experimental deformation measurements were analysed using descriptive statistical methods to evaluate the repeatability and reliability of the ballistic performance. For each ammunition type (9 × 19 mm and 5.56 × 45 mm), three independent impact tests (n = 3) were conducted on intact panels under identical boundary conditions. The maximum back-face deformation (BFD) values were recorded for each shot.

The experimental data were expressed as mean ± standard deviation (SD), calculated according to:

$$SD = \sqrt{{\displaystyle \frac{{\mathsf{\Sigma }(x_i - \overline{x})}^2}{n - 1}}}$$

(1)

where $x_{i }$ represents each individual measurement and $\overline{x}$ the arithmetic mean. Additionally, 95% confidence intervals (CI) were computed assuming normal distribution of experimental data:

$$CI = \overline{X} \pm t_{(0.975,n-1) }\times {\displaystyle \frac{SD}{\sqrt{n}}}$$

(2)

where $t$ corresponds to Student’s t-distribution value for n − 1 degrees of freedom.

Given the limited sample size, inferential statistical comparisons such as ANOVA were not performed. Instead, emphasis was placed on the assessment of experimental variability and its correlation with the numerical predictions. The low standard deviation observed for both calibres indicates acceptable experimental repeatability under controlled testing conditions.

3. Results

3.1. Numerical Simulation Results

Explicit dynamic analyses were performed using ANSYS Workbench, which solves the equations of motion according to the fundamental principles of the finite element method (FEM). The solver employs a central-difference time integration scheme, which is particularly suitable for highly nonlinear and high-velocity events such as ballistic impacts. The governing equilibrium equation solved at each time step is derived from the principle of virtual work [34,35]:

∫_e {δε}^T {σ} dV = ∫_e {δu}^T {F} dV + ∫_s {δu}^T {T} dS

(3)

where {σ} is the vector of Cauchy stresses, {δε} is the virtual deformation tensor, {δu} is the vector of virtual displacements, {F} represents the body forces and {T} represents the applied surface forces.

The material behaviour of the AA5083-H32 aluminum armor plate and the steel projectile was modelled using the Johnson–Cook (JC) constitutive model, which is widely employed to describe plastic deformation under high strain-rate loading conditions. The yield stress is defined as [36]

$$\sigma = [\mathrm{A} + \mathrm{B} {{\epsilon }_{\mathrm{p}}}^{\mathrm{n}}] [1 + \mathrm{C} \ln ({\dot{\epsilon }}_p^{*}\left)\right] [1 - {\mathrm{T}}^{\mathrm{hm}}]$$

(4)

where $A$ is the initial yield strength, $B$ and $n$ are the strain-hardening coefficients, ${\epsilon }_p$ is the equivalent plastic strain, $C$ is the strain-rate sensitivity coefficient, ${\dot{\epsilon }}_p^{*}$ is the dimensionless plastic strain rate, $T^{*}$ is the homologous temperature, and $m$ is the thermal softening coefficient. The Johnson–Cook parameters for AA5083-H32 aluminum were adopted from [36].

The complete Johnson–Cook (JC) parameter set employed in this study is reported in

Table 3. Johnson–Cook constitutive parameters adopted for AA5083.

Material	A (MPa)	B (MPa)	n (-)	C (-)	m (-)	$\dot{\mathsf{\epsilon }}$ 0 (s−1)	T Room (K)	T Melt (K)	D1	D2	D3	D4	D5	Ref.
AA5083	167	596	0.551	0.001	0.859	1.0	293	893	0.0261	0.263	−0.349	0.147	16.8	[37,38]

Note: Parameters reported for AA5083 in the literature were adopted due to the limited availability of JC datasets explicitly calibrated for the H32 temper; this choice is justified because the model is used to reproduce the global ballistic response and is experimentally validated in the present work.

. Although local strain rates during ballistic impact at velocities of approximately 850–920 m·s⁻¹ may reach the order of 10³–10⁵ s⁻¹, the JC model has been widely adopted for AA5083 aluminum in high strain-rate and impact simulations, as it provides robust predictions of the global structural response—including plastic work, residual velocity trends, and back-face deformation—when properly validated.

In the present study, the suitability of this modeling approach is supported by the close agreement between the numerically predicted and experimentally measured maximum deformations, as well as by the observed ductile deformation behaviour of the aluminum layer. Nevertheless, it is acknowledged that the Johnson–Cook model does not explicitly capture microstructural failure mechanisms such as adiabatic shear localisation, interface damage evolution, or delamination in adjacent composite layers. These aspects will be addressed in future work through the implementation of progressive damage models and cohesive-zone formulations.

The model parameters were obtained from [39] and adjusted to the specific conditions of the study, are described in

Table 4. Boundary conditions in the simulation.

Components	Stiffness Behavior	Speed	Support	Thickness
Door—Steel ST37	Flexible	-	Fixed	5 mm
Aluminum 5083-H32	Flexible	-	Fixed	25 mm
Kevlar	Flexible	-	Fixed	15 mm
Cabuya Fiber	Flexible	-	Fixed	15 mm
5.56 × 45 mm ammunition	Rigid	920 m/s	Free	-
9 × 19 mm ammunition	Rigid	360 m/s	Free	-

:

The boundary conditions for the door model were applied considering a rigid fixing system at the edges, which ensures that the projectile’s energy is absorbed mainly by the armor and not by the displacement of the door. It was also assumed that Kevlar and cabuya fibers are isotropic materials in the direction of loading, which is a simplification of the actual anisotropic properties of these materials. This approximation was validated in previous studies for ballistic impact simulations [3].

3.1.1. Reference Simulation (Unshielded Door)

To establish a baseline for comparison, the impact of the projectiles on the original vehicle door structure, modelled as ST37 steel, was simulated. The impact of a 9 × 19 mm projectile resulted in complete penetration of the panel (Figure 7a). A similar behaviour was observed for the 5.56 × 45 mm projectile, which produced catastrophic material rupture in the impact zone (Figure 7b), confirming the inherent vulnerability of the unprotected door structure to ballistic threats.

3.1.2. Armored System Simulation

The implementation of the multilayer Al/Kevlar/cabuya fiber armor system resulted in a significant improvement in structural performance. The numerical model consisted of 17,555 elements and 13,581 nodes, with an average mesh quality of 0.93 according to the Element Quality metric [34], ensuring reliable accuracy of the simulation results.

9 × 19 mm Impact Response: The simulation predicted a maximum localized deformation of 52.75 mm at the point of impact (Figure 8a). The von Mises equivalent stress reached a maximum value of 613.97 MPa, located at the AA5083 aluminum faceplate, indicating substantial energy absorption through plastic deformation. The deformation distribution gradually decreases toward the edges of the panel, suggesting effective dissipation of the projectile’s kinetic energy through plastic work, as described by the constitutive formulation in Equation (4). The rear view of the armor system (Figure 8a) reveals a convex bulge corresponding to back-face deformation (BFD) without visible fracture, demonstrating that the Kevlar and cabuya fiber layers efficiently absorbed the residual impact energy and prevented complete perforation.

5.56 × 45 mm Impact Response: For the 5.56 × 45 mm projectile, the von Mises equivalent stress reached a maximum value of 606.59 MPa, again located in the AA5083 aluminum faceplate (Figure 8b). This result confirms that the aluminum layer underwent significant plastic deformation without reaching full fracture, a behaviour accurately captured by the Johnson–Cook constitutive model (Equation (4)). Importantly, the simulation results show that the stress field was distributed throughout the multilayer armor system (Figure 8c), indicating effective load transfer between the layers. The Kevlar and cabuya fiber composite layers remained structurally intact and continued to contribute to the energy absorption process, preventing penetration of the system.

In all simulated scenarios, the proposed armor system successfully prevented complete perforation, demonstrating its conceptual feasibility as an effective protection solution for vehicular applications. The effectiveness of the design results from the synergistic interaction between the aluminum front layer, which blunts the projectile and absorbs a significant portion of the kinetic energy through plastic deformation described by Equation (4), and the composite layers, which capture fragments and dissipate the remaining energy through tensile deformation mechanisms. All these processes were simulated within the explicit finite element framework governed by Equation (3).

The numerical results (Figure 9) showed stable convergence for both ballistic configurations. For the 5.56 mm projectile, convergence was achieved with variations below 5%, stabilising at a peak von Mises stress value of 606.59 MPa, which demonstrates the robustness of the computational model. Similarly, for the 9 mm projectile, convergence was established at 613.97 MPa, where the stress–time curves exhibited variations lower than the predefined convergence threshold, confirming the reliability of the finite element simulations.

The convergence criterion was defined based on the variation in the peak von Mises stress obtained from successive mesh refinements. Convergence was considered achieved when the variation between consecutive solutions remained below 5%, which is commonly accepted for nonlinear explicit dynamic simulations involving high strain-rate impact events.

3.2. Experimental Validation and Comparative Layer by Layer Analysis

Experimental validation was performed under NIJ 0108.01 and STANAG 4569 Level 1 protocols in a controlled military testing facility. The hybrid armor system successfully resisted both 9 × 19 mm and 5.56 × 45 mm projectile impacts without complete system perforation, thereby satisfying the required ballistic protection criteria.

Global Structural Response: For the 9 × 19 mm projectile, the first aluminum layer underwent significant plastic deformation and partial penetration, while the projectile was ultimately arrested within the internal reinforcement structure. The experimentally measured back-face deformation (BFD) was consistent with the numerically predicted maximum deformation of 52.75 mm, confirming the accuracy of the numerical model in predicting the global structural response. This agreement between numerical predictions and experimental measurements also supports the suitability of the homogenized modeling approach adopted for the composite layers. Although the Kevlar and cabuya fiber composites exhibit anisotropic and strain-rate dependent behaviour, the close correlation between predicted and observed back-face deformation indicates that the simplified formulation is capable of accurately reproducing the dominant global energy dissipation mechanisms governing the ballistic response of the multilayer armor system. This agreement is illustrated in Figure 10, which compares the experimental and numerical deformation results for both projectile calibres.

Under 5.56 × 45 mm impact, the projectile perforated the first aluminum layer and was subsequently arrested within the Kevlar reinforcement stack. The numerical model predicted peak von Mises stresses of 613.97 MPa in the aluminum plate—exceeding its yield strength—which explains the experimentally observed perforation of the frontal metallic layer. In both cases, the close agreement between simulated and experimental deformation patterns, shown in Figure 10, validates the predictive capability of the finite element model and confirms the absence of complete system penetration.

Layer-by-Layer Failure Mechanisms: To extend the validation beyond global performance indicators, a detailed comparison of the observed failure mechanisms was conducted.

Aluminum 5083-H32: The metallic front layer exhibited pronounced ductile plastic deformation, characterised by localized bulging, radial material flow, and the formation of a plastically expanded crater around the impact zone, as shown in Figure 11b–d. In the case of the 9 mm projectile, the aluminum plate showed significant local indentation and material stretching without catastrophic fracture, allowing the projectile to be retained within the structural assembly (Figure 11c). Conversely, under 5.56 mm impact, complete perforation of the front aluminum layer was observed, accompanied by pronounced edge tearing and plastic flow around the penetration site (Figure 11b,d).

The experimentally observed damage morphology correlates with the high stress concentrations predicted numerically in the aluminum layer, confirming that perforation resulted from localized stresses exceeding the material’s yield strength and ultimate tensile limits.

Kevlar 29 (50 plies): The Kevlar layers exhibited a pronounced tensile fibrillation mechanism under 5.56 mm impact, characterised by localized fiber stretching, yarn pull-out, and progressive rupture within the impact zone, as illustrated in Figure 12a. The formation of a deformation cone on the rear face of the laminate indicates effective stress redistribution across multiple plies, allowing the impact energy to be dissipated through inter-fibrillar friction and fiber elongation.

As shown in Figure 12b, the retention of metallic fragments and the projectile core within the Kevlar stack confirms the laminate’s ability to arrest high-velocity threats through progressive tensile failure rather than catastrophic tearing. The experimentally observed fragment capture and conical back-face deformation are consistent with the stress distribution patterns predicted numerically, thereby validating the model’s ability to represent multilayer tensile energy absorption and interlaminar load transfer.

Cabuya Fiber Composite Backing: The treated cabuya fiber layer showed distributed micro-damage without catastrophic delamination. Numerical stress fields in this region confirm its role as a residual energy dissipation and stress redistribution layer.

The comparison presented in

Table 5. Boundary conditions in the simulation. Comparison of numerical predictions and experimental observations.

Layer	Numerical Prediction	Experimental Observation	Correlation Level
Aluminum 5093-H32 (9 mm)	Plastic strain concentration; peak stress 613.97 MPa; no full perforation	Ductile bulging; no complete perforation	High
Aluminum 5083-H32 (5.56 mm)	Localized failure at front plate; stress > 600 MPa	Perforation of first aluminum layers; plastic flow	High
Kevlar 29	Tensile stress redistribution; central strain concentration	Fiber fibrillation; projectile core retained	High
Cabuya Fiber Composite	Distributed stress field; residual deformation	Micro-cracking; no catastrophic delamination	Moderate
Global Response	Maximum deformation 52.75 mm (9 mm)	Measured BFD within predicted range	High

demonstrates a consistent agreement between numerical predictions and experimentally observed damage mechanisms across all structural layers.

From a physical standpoint, the numerical stress distribution predicted in the simulations provides insight into the sequence of failure mechanisms observed experimentally. The concentration of von Mises stresses in the aluminum front layer predicted by the finite element model correlates with the ductile plastic deformation and localized crater formation observed during the ballistics tests. Similarly, the simulated redistribution of stresses toward the composite layers is consistent with the experimentally observed tensile fibrillation and fragment retention within the Kevlar reinforcement. The gradual reduction in stress magnitude toward the rear region of the multilayer system also explains the distributed micro-cracking observed in the treated cabuya fiber composite, confirming its role as a residual energy dissipation layer within the graded armor architecture.

To contextualize the ballistic response of the proposed multilayer vehicular armor, performance was expressed using normalized metrics, namely areal density (AD), absorbed energy per unit areal density (Eabs/AD), specific energy absorption (SEA), and back-face deformation (BFD). These parameters are widely employed in ballistic composite research to enable quantitative comparison across distinct material systems and architectural configurations, independent of variations in geometry or test setup. Representative composite armor systems reported in the literature—including UHMWPE/aramid hybrids, ceramic–fiber multilayers, and three-dimensional woven architectures—are summarized in

Table 6. Comparative Ballistic Performance of Composite Armor Systems.

System/Architecture	Threat	AD (kg/m2)	Eabs/AD (J·m2/kg)	SEA (J/kg)	BFD/BFS	Reference
5083-H32/Kevlar 29/cabuya fiber	9 × 19 and 5.56 × 45, no perforation	140.87	3.68 (9 mm)/10.70 (5.56 mm)	4.94 (9 mm)/ 14.35 (5.56 mm)	52.75 mm	-
UHMWPE/modified Aramid	7.62 × 25 mm (soft armor)	42.0	147.2	1635.0	19	[40]
UHMWPE UD sheets	9 mm	4.72	153.6	960.3	9.86	[41]
3D Kevlar/UHMWPE	NIJ Std soft armor	5.94	91.7	573.0	19.6	[42]
Ceramic/Aramid/UHMWPE	12.7 mm API	3.42	175.8	1098.8	28.65	[43]

Normalized metrics: Areal Density (AD), energy (Eabs) normalized by AD (Eabs/AD) and Specific Energy Absorption (SEA).

.

3.3. Post-Impact Microstructural Damage Analysis

Figure 13 presents the progressive microstructural damage observed in the 5083-H32 aluminum front layer. The unimpacted material (Figure 13a) exhibits a homogeneous surface morphology without evidence of pre-existing microcracks or plastic deformation, serving as a reference condition.

Following ballistic impact, a pronounced radial plastic wave pattern becomes evident (Figure 13b), characterized by concentric deformation bands consistent with high strain-rate plastic flow. This morphology reflects the propagation of compressive shock waves and shear-dominated plastic deformation, confirming that the aluminum layer acts as the primary energy dissipation stage through dynamic yielding.

At higher magnification (Figure 13c), localized material loss and micro-extrusion are observed in the impact region, indicating the onset of adiabatic shear localization and crater formation. These features correlate strongly with the elevated von Mises stresses (~614 MPa) predicted by the Johnson–Cook constitutive model. The absence of brittle fracture supports the assumption that plastic deformation governs projectile deceleration in the initial impact phase.

Overall, the experimentally observed progressive damage morphology validates the numerical prediction of strain-rate dependent plastic flow as the dominant energy absorption mechanism in the metallic front layer.

Figure 14 illustrates the progressive microstructural damage evolution in the Kevlar 29 intermediate layer. The unimpacted reference state (Figure 14a) shows continuous and well-aligned aramid filaments, characteristic of the woven architecture employed.

Following ballistic impact, significant fibrillation and longitudinal splitting are observed (Figure 14b). The separation of microfilaments and increased interfilament spacing indicate tensile-dominated deformation accompanied by frictional energy dissipation between fibrils. This mechanism is widely recognized as a primary energy absorption process in aramid-based ballistic systems.

At higher damage levels (Figure 14c), localized fiber rupture is evident in regions of maximum tensile concentration. Importantly, fracture is not uniformly transverse but rather confined to localized zones, suggesting progressive load redistribution rather than catastrophic ply failure. This observation aligns with the principal stress distribution predicted numerically, where residual stresses transferred from the plastically deformed aluminum layer concentrate within the Kevlar reinforcement.

The combined fibrillation, fiber stretching, and localized rupture confirm that the Kevlar layer functions as a high-tensile energy absorber, operating through progressive axial deformation rather than brittle fragmentation.

Figure 15 presents the microstructural condition of the treated cabuya fiber composite before and after ballistic loading. In the unimpacted state (Figure 15a), the fiber–matrix interface appears continuous and well bonded, with clear epoxy impregnation along the roughened fiber surface produced by the alkaline treatment. No interfacial voids or debonding zones are observed, indicating effective load transfer capability within the composite backing. Following ballistic impact (Figure 15b), partial fiber fragmentation and localized longitudinal splitting are evident. Importantly, the failure is not catastrophic; instead, fiber rupture occurs progressively, accompanied by limited pull-out and micro-fracture of individual filaments. The absence of widespread interfacial delamination suggests that energy dissipation occurs through controlled fiber breakage and frictional interaction within the matrix.

These observations support the hypothesis that the treated cabuya layer acts as a residual stress redistribution medium rather than a brittle structural element. The progressive fragmentation mechanism correlates with the numerically predicted reduction in stress magnitude toward the rear surface, confirming its role as the final stage in the multi-layer energy dissipation sequence.

4. Discussion

This study demonstrates the successful development and validation of a hybrid multilayer armor system integrating aluminum, Kevlar, and cabuya fiber for vehicular protection. The results provide relevant insights into the advancement of ballistic protection materials, particularly within multilayer hybrid architectures. Recent studies have investigated four-layer armor systems reinforced with carbon nanotubes CNTs 37, reporting significant improvements in energy absorption and ballistic resistance. These findings highlight the importance of optimizing material combinations, consistent with the hybrid design strategy proposed in the present work, where high-performance materials such as Kevlar and natural fibers are combined with metallic layers to maximize protective performance.

Compared with conventional single-material solutions, material hybridization enables more efficient dissipation of kinetic energy, as demonstrated in several previous studies [12,20]. The proposed system not only enhances energy absorption efficiency but also reduces structural weight, which is a critical requirement for vehicular armor applications [44].

The previous literature [8,45] has emphasized that the ballistic performance of composite materials such as Kevlar and natural fibers strongly depends on projectile velocity and ammunition type. According to [7], Kevlar dissipates energy primarily through fibrillation mechanisms, which play a key role in the performance of hybrid armor systems. This behaviour was confirmed in the present experimental tests, where the 5.56 mm projectile core was retained within the Kevlar layer, consistent with the predictions obtained from the numerical simulations.

The incorporation of natural fibers, beyond their mechanical contribution, also enhances the sustainability profile of the system. The use of biodegradable fibers in protective structures has been widely discussed in the literature [20,46]. Materials such as sisal and cabuya offer not only reduced production costs but also more environmentally sustainable alternatives in a sector traditionally dominated by synthetic fibers [6].

The comparative analysis presented in Table 5 highlights distinct ballistic performance trends among different composite armor architectures. Lightweight UHMWPE and modified aramid systems subjected to 7.62 × 25 mm and 9 mm threats demonstrate significantly higher specific energy absorption (SEA), reaching values above 960–1635 J/kg [40,41]. This elevated efficiency is mainly attributed to the tensile-dominated energy dissipation mechanisms characteristic of ultra-high-molecular-weight polyethylene laminates, which maximize fibrillation and membrane stretching during impact.

The strong correspondence between numerical stress fields and experimentally observed failure patterns confirms that the finite element model successfully captures the dominant energy dissipation mechanisms governing the ballistic response of the hybrid multilayer system. In particular, the numerical model reproduces the sequential interaction between layers, where the aluminum plate undergoes plastic deformation and projectile blunting, followed by tensile energy absorption within the Kevlar reinforcement and final stress redistribution in the natural fiber composite backing [11]. This agreement indicates that the explicit dynamic modeling approach is capable of representing the global failure sequence of the multilayer armor system, even though certain local composite damage mechanisms remain simplified within the homogenized formulation.

Similarly, three-dimensional Kevlar/UHMWPE woven architectures exhibit improved control of back-face deformation (BFD ≈ 19.6 mm) due to enhanced through-thickness reinforcement and increased inter-yarn friction [42]. Ceramic–aramid–UHMWPE hybrid configurations subjected to higher-energy 12.7 mm API threats exhibit superior Eabs/AD ratios (≈175.8 J·m²/kg), primarily governed by brittle ceramic fracture and localized projectile erosion mechanisms [43].

In contrast, the aluminum–Kevlar–cabuya fiber system proposed in this study follows a structurally integrated design philosophy. Although its SEA values (4.94–14.35 J/kg) are lower than those reported for textile-dominated armor systems, the architecture promotes sequential energy redistribution through metallic plastic deformation, tensile energy absorption in Kevlar, and controlled damage within the cabuya fiber composite backing. This multistage dissipation mechanism, combined with a door-level structural configuration (AD = 140.87 kg/m²), differentiates the proposed system from flat soft-armor panels and positions it within the context of vehicular structural reinforcement rather than lightweight personal protection systems.

Weight reduction remains a critical design objective in the development of efficient vehicular armor systems, as consistently emphasized in recent ballistic materials research [47,48]. In this context, the partial replacement of expensive synthetic reinforcements with natural fibers such as cabuya contributes not only to mass optimization but also to measurable economic and environmental benefits [10]. Market data reported in the literature indicate that aramid fibers typically cost approximately 35–60 USD/kg, whereas natural lignocellulosic fibers are commonly available in the range of 1–3 USD/kg, depending on fiber type and processing conditions [49]. Furthermore, the embodied energy associated with aramid fiber production has been reported to range between approximately 150–200 MJ/kg, whereas natural fibers typically exhibit much lower embodied energy values, generally within 5–15 MJ/kg, reflecting their lower processing requirements and renewable origin [12,50]. These values should be interpreted as approximate ranges reported in the literature rather than precise cost metrics, as they may vary depending on geographic availability, processing methods, and industrial scale.

In addition to hybrid configurations, aramid-only laminate systems have been extensively investigated as baseline ballistic protection solutions. Recent studies on Kevlar and Twaron panels subjected to NIJ-compliant 9 mm FMJ impacts report that their energy dissipation is predominantly governed by tensile membrane stretching, yarn fibrillation, and progressive fiber rupture within the textile plies, with damage localized mainly in the impact zone [51,52]. Although such systems achieve relatively high specific energy absorption due to their low density and efficient tensile response, their performance relies primarily on in-plane fiber engagement without graded stress redistribution through heterogeneous layers. In contrast, the aluminum–Kevlar–cabuya architecture proposed in this study introduces an additional dissipation stage following Kevlar engagement, where the treated natural fiber composite backing promotes controlled micro-cracking and attenuation of residual stress waves. This mechanism modifies the post-impact load-transfer pathway and enhances structural-scale deformation control, demonstrating that the natural fiber layer functions as an active energy redistribution component rather than a passive mass-reduction filler [51,52].

Overall, compared with other hybrid armor configurations reported in the literature, the proposed aluminum–Kevlar–cabuya system offers a distinct balance between structural integrity, weight reduction, and certified ballistic resistance at door scale. Its main advantage lies in the integration of a natural fiber composite as an active residual energy redistribution layer within a structural vehicular protection architecture. This approach contributes to reducing both material cost and environmental impact while maintaining non-perforation under 9 × 19 mm and 5.56 × 45 mm FMJ threats However, in contrast to lightweight textile-dominated armor systems which typically exhibit higher specific energy absorption due to their lower areal density—the present configuration is characterized by a higher areal density and lower SEA values, reflecting its intended role as a structural vehicular armor solution rather than a personal protection panel. Although the system successfully prevented complete penetration, the measured back-face deformation remains larger than that reported for some advanced textile- and ceramic-based hybrid configurations. Therefore, the proposed architecture demonstrates clear advantages in terms of structural-scale applicability, cost efficiency, and multistage energy dissipation, while its main limitation remains the relatively high deformation amplitude associated with high-energy impact events. Despite the strong agreement between numerical predictions and experimental observations at the structural scale, certain modeling limitations must be acknowledged. From a modeling perspective, the use of an equivalent homogenised formulation primarily affects the prediction of local damage mechanisms within the composite layers rather than the overall structural behaviour of the armor system. Parameters such as penetration resistance, global deformation patterns, stress redistribution between layers, and back-face deformation are governed primarily by the multilayer structure’s global stiffness and energy-absorption capacity. Therefore, although the model does not resolve detailed fiber-level failure mechanisms, it remains suitable for evaluating the global ballistic response of the hybrid armor architecture. In the present work, the Kevlar and cabuya composite layers were represented using an equivalent homogenized formulation derived from orthotropic elastic data. This approach provides stable and computationally efficient prediction of the global ballistic response, including deformation patterns, stress redistribution, and penetration resistance. However, the model does not explicitly incorporate strain-rate-dependent orthotropic failure criteria, such as Hashin, Puck, or continuum damage mechanics formulations. Consequently, localized intra-ply failure mechanisms—including progressive fiber rupture, yarn pull-out, and interlaminar delamination—cannot be resolved with full fidelity. While this limitation does not significantly affect predictions of global structural performance, it may constrain the accuracy of local damage evolution within the composite layers. To further enhance performance, future research may explore the integration of polyurea coatings, which have demonstrated the ability to restore degraded armor performance and reduce back-face deformation (BFD) through improved energy dissipation and crack-bridging mechanisms [53,54], Such hybridization strategies could mitigate the primary limitation observed in this study—namely the magnitude of deformation under high-energy impacts—while preserving the cost-efficiency and sustainability benefits of the proposed architecture [55,56].

5. Conclusions

This study validated the hypothesis that incorporating a treated cabuya fiber composite as a structural backing layer enhances multi-stage energy dissipation in a graded aluminum–aramid hybrid armor system without compromising certified ballistic resistance.

The multilayer configuration (25 mm AA5083-H32/15 mm Kevlar 29/15 mm cabuya composite) successfully resisted 9 × 19 mm and 5.56 × 45 mm FMJ threats under NIJ 0108.01 and STANAG 4569 Level 1 conditions, preventing complete perforation. Numerical predictions indicated a maximum back-face deformation of 52.75 mm under 9 mm impact, showing strong agreement with the experimental measurements (mean ± SD, n = 3). Peak von Mises stresses (~614 MPa) in the aluminum layer explain the observed localized plastic deformation during rifle impact while maintaining structural integrity.

Layer-by-layer analysis confirmed a sequential energy dissipation mechanism governed by:

• strain-rate-dependent plastic deformation and projectile blunting in the aluminum front layer;
• tensile fibrillation and fragment retention within the Kevlar plies;
• residual stress redistribution and controlled micro-cracking in the treated cabuya composite backing.

With an areal density of 140.87 kg/m² and a total door mass of 105 kg, the system achieved a 19% weight reduction compared with conventional steel-based armouring while maintaining structural-scale ballistic resistance.

The main novelty of this work lies in the experimentally validated integration of treated cabuya fiber composite as an active structural energy redistribution layer within a certified aluminum–aramid architecture at door scale, supported by an explicit dynamic finite element framework capable of predicting sequential multi-material dissipation mechanisms.

Although the numerical framework successfully reproduces the global ballistic response of the hybrid structure, the composite layers were represented using an equivalent homogenized formulation. Consequently, strain-rate-dependent orthotropic failure mechanisms—such as progressive fiber rupture, yarn pull-out, and interlaminar delamination—are not fully resolved in the current model. This limitation mainly affects the prediction of localized intra-ply damage, rather than the global structural response of the armor system.

Future research will focus on incorporating strain-rate-dependent orthotropic damage models and optimizing layer topology and thickness distribution, and evaluating system performance against higher kinetic energy threats (example 7.62 × 51 mm). The validated configuration demonstrates practical potential for integration in civilian high-risk vehicles, security fleets, and lightweight armoured transport platforms requiring cost-effective and sustainable ballistic protection solutions.