Research on Basic Properties of Polymers for Fused Deposition Modelling Technology

Abstract

This study investigates the mechanical properties and biocompatibility of eight commercially available filaments tailored for Fused Deposition Modeling (FDM) additive manufacturing. Test specimens were fabricated using original PRUSA MK4 printers, with ten samples from each selected polymer. Mechanical evaluations through static tensile and three-point bending tests revealed that PETG Carbon and PA+15CF exhibited superior tensile and flexural strengths, making them highly suitable for applications requiring high mechanical resilience. Biocompatibility assessments in line with the ISO 10993-5:2009 and ISO 10993-12:2021 standards indicated that all materials except FiberFlex 40D Fiberlogy were non-cytotoxic, supporting their potential in biomedical applications. The experimental data established material constants within the Johnson–Cook strength model, which effectively predicted the mechanical behaviors of monotonic materials like FiberFlex 40D, PETG, HIPS, TPU, and PA+15CF Rosa 3D, with maximum fitting errors not exceeding 2.6%. However, the model was inadequate for non-monotonic materials like PLA and PETG, resulting in higher errors and less accurate simulations. Scanning electron microscope (SEM) analyses provided insights into fracture mechanisms, correlating fracture surface characteristics with mechanical performance. This comprehensive study advances the understanding of mechanical properties in thermoplastic materials for 3D printing, validates numerical models for certain materials, and confirms material suitability for biomedical use.

1. Introduction

Three-dimensional (3D) printing technologies using Fused Filament Fabrication (FFF) and Fused Deposition Modeling (FDM) have become among the most prevalent methods of additive manufacturing for small- and medium-scale production. This popularity is largely attributed to the relatively simple design and operational ease of FDM printing devices, as well as the accessible costs of both the printers and thermoplastic materials. These factors have facilitated the widespread adoption of this technology and fostered a large and active community of users. Such accessibility has allowed 3D printing to become a versatile tool within engineering, with applications spanning from industrial prototyping to highly specialized medical applications. In particular, the role of 3D printing in rapid prototyping has proven invaluable for medicine, including fields such as orthopedics, surgery, cardiology, and prosthetics [1,2,3,4,5,6,7,8,9]. The capacity to develop a detailed, three-dimensional model of an anatomical structure for preoperative planning, for instance, is crucial in many medical contexts [10,11]. Additionally, objects fabricated with FDM that are intended for external use, such as prostheses, rehabilitation devices, and medical equipment, can be customized in ways that significantly enhance their functionality and fit.

Moreover, the flexibility of FDM technology to incorporate various materials—ranging from carbon fiber-reinforced polymers to flexible filaments such as rubbers and silicones—broadens its applicability. The precision of the FDM technique is often sufficient for external medical applications, and the technology’s ease of manufacture, low cost, and availability allow for the rapid production and distribution of these elements, particularly in scenarios where access to medical facilities may be limited or where a temporary solution is necessary. Despite these strengths, FDM technology also has its limitations. A significant drawback is the restricted range of compatible materials, which can limit the applicability of FDM for biomedical implants and other internal applications. While external applications may tolerate the current level of precision, internal applications often require additional post-processing steps to meet the stringent accuracy and reliability standards expected within the body. Nonetheless, the technology is advancing rapidly, and recent developments show promise in addressing these limitations, especially through enhancements in material properties and printing precision [12,13,14].

Given the evolving landscape of FDM technology, it is crucial to evaluate both the mechanical and biological properties of the materials used in these additive manufacturing processes. Over the past decade, extensive studies have been conducted to assess the impact of the FDM process on material properties, as well as the behavior of various filaments under different conditions [15,16,17,18,19,20,21,22,23]. Valean [24], for example, demonstrated how variables such as the printing orientation and material thickness impact mechanical properties, identifying the optimal printing orientation for enhanced predictability in material behavior during strength testing. Similarly, Tymrak [25] examined the basic mechanical characteristics of PLA and ABS filaments, yielding valuable insights into their tensile strength and elastic modulus. However, while such studies have laid a foundation, they primarily focus on widely used filaments like PLA and ABS, overlooking more complex materials, particularly those with composite structures, such as polyamides reinforced with carbon fibers. Current trends are encouraging attempts to use additive technologies for biomedical purposes in an attempt to study both cytotoxicity of materials or attempt to use new materials adapted to the human body [26]. Addressing this gap, Grygier et al. [15] conducted pioneering work in examining a set of related materials within FDM technology, establishing a framework for the systematic study of diverse filament types.

Building upon these prior studies, the present research aims to expand the understanding of FDM materials by investigating eight commercially available filaments, including composite and specialized types, to assess their mechanical properties, fracture mechanisms, and biocompatibility. Specifically, this study employs tensile [27] and three-point bending testing, SEM analysis of fracture surfaces, and biocompatibility testing aligned with ISO 10993-5 [28] standards. The outcomes of this research will not only contribute to filling the identified gaps in the literature by providing comprehensive data on a broader range of FDM materials but will also support the development of more accurate computational model libraries for these materials. By addressing these gaps, the findings of this study hold significant implications for biomedical and industrial applications, underscoring the growing relevance and adaptability of FDM technology within the academic and engineering communities.

2. Materials and Methods

2.1. Characteristics of Printed Materials

The research included a set of materials including thermoplastics intended for additive printing using FDM technology. FDM (Fused Deposition Modeling) is a method that involves applying a thermoplastic material layer by layer, after which the filament is heated in the so-called extruder and is then applied in layers using a head. These materials differed in terms of their physical parameters. Each sample had 5 upper and lower compact layers, 5 outlines, and 100% fill. The printing speed was unified for each type of sample filling and was 30 mm/s. The layer height was 0.2 mm. The extrusion temperature of each filament was selected individually according to the material properties.

Table 1. Materials used in the study [29,30,31,32,33,34,35].

No.	Material	Manufacturer	Density (g/cm3)	Tensile Strength Declared by Producent (MPa)	Hardness (HS)	Volumetric Shrinkage (%)
1	PETG Carbon	Kimya	1.32	92.9	78.8	0.2–1.0
2	FiberFlex 40D	Fiberlogy	1.27	26.5	40.0	0.5–1.0
3	PETG	Sunlu	1.23	28.3–101.0	71.4	0.2–1.0
4	HIPS	Devil Design	1.05	37.0	75.0	0.2–1.0
5	TPU	F3D Filament	1.20	28.0–96.0	85.0	0.5–1.0
6	PA+15CF	ROSA 3D	1.07	125.0	78.0	0.1–0.5
7	PETG	3DActive	1.26	84.6	70.0	0.2–1.0
8	PLA	3DActive	1.24	45.0	87.0	0.3–0.5

contains detailed information on the selected types of filaments, along with data on their mechanical properties as provided by the manufacturers. Material samples were prepared for printing using the original Prusa MKIV printer (Prague, Czech Republic, PrusaResearch).

The materials used in the study were selected by distinguishing their properties. Materials such as FiberFlex and TPU exhibit flexible behavior, while PLA, PETG, HIPS, and PA exhibit brittle properties. By examining their properties and then comparing them, it is possible to estimate the appropriate application for each of these materials. Materials with carbon fiber additives demonstrated high stiffness, with optimal results achieved under carefully controlled extrusion and bed temperatures to address surface roughness and material detachment issues [36,37]. Flexible filaments posed challenges such as nozzle clogging and excessive flow; however, reducing the extrusion temperatures improved smoothness and dimensional accuracy [29]. Thermoplastics like HIPS required specific adjustments to counteract material contractions and brittleness, with modifications such as increased bed temperatures and fan cooling adjustments yielding consistent, smooth prints [30]. TPU, characterized by abrasion resistance and low shrinkage, provided highly precise and flexible prints under recommended conditions, while polyamide-based composites reinforced with carbon fibers offered uniform layer distribution when printed at high temperatures [31]. Widely used materials like PLA and PETG required relatively standard conditions, producing consistent results with smooth surfaces and no dimensional deviations. The findings underline the importance of optimizing printing parameters for each material to achieve desirable properties, contributing valuable insights into enhancing FDM print performance.

Before fabricating the final test specimens, preliminary experiments were conducted to determine the optimal printing temperatures for extrusion and bed heating, both of which play a crucial role in ensuring print quality and mechanical performance. The optimization process involved producing calibration samples with dimensions of 20 × 20 mm and a height of 15 mm (Figure 1). These calibration prints were systematically evaluated for their surface finish, dimensional accuracy, and structural integrity to identify the temperature settings that yielded the best extrusion parameters. For extrusion temperatures, the focus was on eliminating common printing defects such as stringing, delamination between layers, and deformation caused by detachment from the heated bed. The analysis involved observing the material’s flow consistency, layer adhesion, and overall structural quality. The optimal extrusion temperature was defined as the temperature at which the filament exhibited smooth flow, strong interlayer bonding, and minimal visual defects. Bed heating temperatures were equally critical to ensure proper adhesion between the printed layers and the printing surface. Inadequate bed adhesion can lead to warping, curling, or layer misalignment, compromising the quality and dimensional accuracy of the specimens. For each type of filament, the bed temperatures were fine-tuned to achieve uniform adhesion and prevent lifting at the edges of the prints, a common issue in materials like PETG and flexible filaments.

The adjustment process revealed that most filaments performed well within the temperature ranges recommended by their manufacturers. However, exceptions were noted for FiberFlex 40D and PETG. For these materials, the manufacturer did not provide precise guidelines for optimal extrusion temperatures, necessitating an experimental approach by the researchers. Through systematic testing and evaluation, the researchers established custom temperature settings that produced defect-free prints with consistent mechanical properties.

The selected temperatures used to produce the test samples are presented in the table below (

Table 2. Temperatures of materials used in the study [29,30,31,32,33,34,35].

No.	Material	Producent	Temperature Recommended by the Manufacturer (°C)	Selected Temperature Printing (°C)	Bed Temperature (°C)
1	PETG Carbon	Kimya	225	225	90
2	FiberFlex 40D	Fiberlogy	220	240	60
3	PETG	Sunlu	250	240	90
4	HIPS	Devil Design	240	240	110
5	TPU	F3D Filament	230	230	60
6	PA+15CF	ROSA 3D	290	290	110
7	PETG	3DActive	250	250	90
8	PLA	3DActive	220	220	60

).

2.2. Fabrication of the Samples

The test specimens were designed and created in accordance with ISO 527-3:2019-01 [27] for tensile testing of polymers. For three-point bending tests, notched samples were used, which are critical for evaluating the material’s resistance to fracture and its behavior under complex stress states. The presence of a notch helps focus the stress on a specific point, providing insights into the material’s fracture mechanics.

All samples were manufactured using original Prusa MK4 3D printers. During the printing process, the specimens were positioned flat on their largest surface to ensure continuous material deposition along the direction of the test. This alignment minimized layer discontinuities and improved the comparability of results by aligning the printed layers with the expected stress direction during testing. This approach also reduced the risk of delamination or interlayer cracking, which is common in FDM-printed parts, thus improving the overall reliability of the results.

Each sample had 5 upper and lower compact layers, 5 outlines, and 100% fill (Figure 2). The printing speed was unified for each type of sample filling and was 30 mm/s. The layer height was 0.2 mm. The extrusion temperature of each filament was selected individually according to the material properties, as stated previously.

2.3. Testing of Materials Using a Static Tensile Test

In accordance with ISO 527-3:2019-01 [27] and ISO 527-2:2012 [38], specific conditions and procedures for tensile testing of plastics were adopted. These standards define the shape and dimensions of the type 1B specimen, including a total length of 150 mm, a distance between shoulders of 115 mm, a gauge width of 10 mm, and a thickness of 4 mm with a tolerance of ±0.2 mm, as shown in Figure 3. The selection of the type 1B specimen aligns with ISO 527-3:2019-01, which recommends this type for rigid materials, such as rigid sheets, where smaller dimensions improve the measurement accuracy. This choice is particularly relevant for FDM-produced samples, which often exhibit layered and stiff structures. The compact dimensions of type 1B enhance precision by reducing deformation effects, thus ensuring the reliable assessment of mechanical properties in the context of additive manufacturing.

The tensile tests were conducted under controlled environmental conditions to ensure compliance with ISO standards. In accordance with ISO 527-1:2019 [39] and ISO 527-2:2012 [38], the testing environment was maintained at a temperature of 23 °C and a relative humidity of 50%.

The tensile test procedure involved gradually stretching carefully prepared flat specimens at a constant speed. The test was performed on a Bionix MTS testing machine (Figure 4) equipped with specialized jaws for tensile testing, a dynamometer for measuring the applied force, and a displacement sensor to capture the elongation relative to the specimen’s initial length.

In accordance with ISO 527-2:2012 [38] and ISO 527-1:2019 [39], different speeds were applied depending on the measurement objective. To determine Young’s modulus, a jaw speed of 1 mm/min was used, ensuring precise strain measurements, which are necessary for modulus calculation, as per the ISO 527-2:2012 [38] recommendations for rigid and semi-rigid plastics. For the tensile test conducted until rupture, a higher speed of 5 mm/min was applied. The testing machine was configured with a maximum jaw separation of 50 mm, corresponding to the maximum allowable elongation. Each sample was placed symmetrically within the machine grips to ensure even force distribution. One tensile test was conducted per sample, with the test concluding when the sample reached the point of fracture.

2.4. Testing of Materials Using a Static Three-Point Bending Test

The principles and procedure for carrying out bending tests of plastics are specified in the ISO 178 standard [37]. This method involves evenly distributing the force in the central part of the sample between the supports. A cuboid-shaped sample with an indentation indicating the expected location of the crack was used to carry out the test (Figure 5).

The bend test procedure involves placing a material sample between two supports and then applying a force to the center of the sample using a third element. This force is gradually increased, causing the sample to bend until it breaks. During the test, both the force and deformation of the sample are recorded, allowing for the analysis of material characteristics such as the strength and elasticity in response to bending. An analysis of the recorded force and deformation data makes it possible to determine the bending resistance characteristics of the tested plastic samples, which provide information about their behavior under load and their mechanical properties. The test was carried out similarly to the static tensile test on the MTS Bionix testing machine (MTS Systems, Eden Prairie, MN, USA) (Figure 6).

2.5. Methodology of Tensile Test Numerical Simulation

Numerical methods are currently very popular in solving engineering problems. They represent a branch of mathematics focused on the creation, design, and analysis of algorithms to address the complex mathematical challenges encountered in physical, engineering, and medical applications. These methods use functions and equations to derive approximate solutions, as the problems analyzed often lack exact analytical solutions. The core objective in numerical modeling is to simplify complex issues to approximate real solutions as closely as possible.

In this study, eight material models of filaments were developed to enhance and expand existing material models. The numerical model of the sample was created based on the actual samples used in static tensile testing (Figure 3). However, it is important to note certain limitations of this study. The Johnson–Cook model was simplified by omitting the strain rate and temperature effects, which may impact the material’s response under varying loading conditions. The parameters for the Johnson–Cook model were determined through an iterative optimization process, minimizing the differences between the fitted curve and the experimentally obtained stress–strain data. The derived parameters were then implemented in numerical simulations, and the resulting curves were compared with the experimental data to validate the model’s accuracy and applicability. Furthermore, the finite element mesh was designed with element sizes of 1 to 2 mm to balance accuracy and computational efficiency. Although a mesh sensitivity study confirmed that the chosen discretization did not significantly affect the results, minor influences on the stress concentration accuracy in high-strain areas could still be present. Lastly, simplified boundary conditions were applied to represent clamping, which may not fully replicate the real interaction between the sample and the testing machine grips.

The sample was discretized with hex elements from the dynamic explicit analysis library in the ABAQUS numerical analysis program. The size of the elements was from 2 mm where they were fixed in the jaws of the testing machine to 1 mm where they were stretched, which resulted in a total number of finite elements of 10,880. The discretized sample is shown below (Figure 7).

The initial boundary conditions were based on the sample clamping conditions. In the areas marked below (Figure 8), the appropriate degrees of freedom were constrained. On one surface, all translational and rotational degrees of freedom were removed, while on the opposite, tensioned surface, all degrees of freedom except translation along the X-axis were removed. Subsequently, an initial velocity of 5 mm/min along the X-axis was applied to the sample.

In order to describe the strength limit of materials, constitutive strength models are commonly used. One of the most important currently used equations representing the strengthening is the Johnson–Cook model, which is a special type of Mises plasticity model, with analytical forms of the law of material hardening and softening under temperature influence. The J-C constitutive equation is suitable for computational purposes due to its simplicity and thanks to the use of parameters that are relatively easy to determine. It is commonly used in CAE programs, such as ABAQUS and LS-Dyna. The equation takes into account the influence of the strain magnitude, strain rate, and temperature on the value of the plastic flow limit. At the current stage of research, the authors did not focus on the influence of the strain rate or temperature, so in a form usable in the context of the current research, the formula looks as follows (1):

$${\sigma }_y=(A+B{\underset{\_}{\epsilon }}^{p^n})$$

(1)

where A is the yield strength [MPa], B is the strengthening constant [MPa], n is the strengthening exponent [-], and ${\underset{\_}{\epsilon }}^p$ is the effective plastic strain [-].

Based on previously conducted tests—the static tensile test—appropriate material parameters were determined for each filament material, which are listed in the table below (

Table 3. Obtained filament material parameters based on experimental research.

Material	E [MPa]	v [-]	ρ [kg/m3]	A [MPa]	B [MPa]	n [-]	Failure Strain [mm/mm]
PETG Carbon	2071	0.4	1320	43.67	25.14	0.099	0.057
FiberFlex 40D	14	0.4	1270	2.43	0.64	1.444	5.295
PETG	1130	0.4	1230	-	-	-	0.030
HIPS	646	0.4	1050	10.80	-	-	0.375
TPU	26	0.4	1200	2.03	1.63	1.723	2.860
PA+15CF ROSA 3D	1667	0.4	1070	30.04	90.17	0.392	0.082
PET-G	970	0.4	1260	33.57	17.40	0.262	0.105
PLA	1488	0.4	1240	40.16	16.55	0.374	0.103

).

A model based on the hydrodynamics of SPH particles was adopted as the failure criterion. In this application, it involves specifying the value of the total strain at which finite elements reaching the value of this strain are converted into SPH particles. Thanks to this, it is possible to observe the propagation of damage with one input value, which reduces the need to conduct a number of different tests to determine the parameter values in various damage models. The values of these strains for individual filament material models are presented in the table above (Table 3). After completing the process of preparing numerical models, a series of numerical simulations were performed for each material.

2.6. Methodology of Three-Point Bending Test Numerical Simulation

First, based on the actual dimensions of the samples, a geometric model of the sample was developed in a computational environment. Additionally, geometric models of the supports and the element loading the sample were also prepared.

All elements were discretized with 1 mm elements from the Abaqus/Explicit dynamic analysis library. For the sample, the total number of elements was 16,850, and for the supports and the loading element, it was 1200 each. The discretized elements of the system are shown below (Figure 9).

The last element was the application of initial boundary conditions. All degrees of freedom were removed from the supports, and the speed of the loading element was set to 5 mm/min. General contact with a friction coefficient of 0.3 was used. The sample with initial boundary conditions is shown below (Figure 10).

In order to mathematically describe the behavior of the material, the Johnson–Cook constitutive model described in the previous section was also used. This also allows the developed mathematical models to be verified in other load variants apart from the static tensile test. Ordinary structural steel in the elastic range was used as the material for the supports and the loading element.

2.7. Methodology of SEM Observations

Observations of the condition of the crack surface after carrying out strength tests in accordance with the ISO 527-3:2019 [27] standard for a static tensile test were made using a HITACHI TM-3000 scanning microscope (Angstrom Scientific, Inc., Tokyo, Japan), equipped with an EDS/EDX detector (EDAX, Pleasanton, CA, USA). Tests were performed using a secondary electron (SE) detector at an accelerating voltage of 15 kV in compo mode (composition) and using a module for observing the topography of the topo surface. Observations were made on samples coated with sputtered graphite, which allowed for a detailed analysis of the fracture surface.

2.8. Biological Assessment of Samples

Medical devices must be biocompatible in terms of the ISO 10993-5 [28] and ISO 10993-12 [36] series of standards, appropriately taking into account their intended purpose and biological classification of body contact, as specified in ISO 10993-1:2018 [40]. For medical devices intended for non-European product registrations, additional requirements, such as the FDA Guidance document “Use of International Standard ISO 10993-1, Biological Evaluation of Medical Devices Part 1: Evaluation and Testing within a Risk Management Process”, dated 4 September 2020, must be taken into account. As requested in ISO 10993-1, the actual biocompatibility test program to be executed for a particular medical device must be based upon risk management considerations, as described in ISO 14971 [41].

Cytotoxicity tests employing cell culture techniques can be used to determine cell death (e.g., cell lysis), cell growth inhibition, colony formation, and other effects on cells caused by medical devices, materials, and/or their extracts. If testing is performed, it shall be conducted in accordance with ISO 10993-5 [28]. The purpose of the study is to determine the potential biological reactivity of mammalian cell culture (L929) in response to the test article extract. The test article is extracted in Minimum Essential Medium (MEM) with 10% fetal bovine serum for 24 h at 37 °C or 72 h at 37 °C for permanent implants or prolonged exposure medical devices. The cells are allowed to grow to sub-confluency in tissue culture plates and are exposed to the neat (100%) test article extract in 3 replicates. The plates are incubated for 48 h at 37 °C. The biological reactivity of the cells following exposure to the extracts is visually observed with a microscope and graded on a scale of 0 (no reactivity) to 4 (severe reactivity). The test article is considered non-cytotoxic if none of the cultures exposed to the test article extract shows greater than mild reactivity (grade 2).

The scope of the study included the testing of the cytotoxic properties of the substance carried out in accordance with the principles of the PN-EN ISO 10993-5:2009 [28] and ISO 10993-12:2021 [36] standards. The evaluation of the cytotoxic potential with the L929 cell line using an MEM elution assay is described by the International Organization for Standardization, ISO 10993-5:2009 [28]. Before the study, the material was subjected to EO sterilization. The parameters of sterilization were 55 °C and a 12 h EO cycle including 3 min post-cycle aeration. Sterilization was conducted (STERI-VAG 5XL—3M, Solventum, North Ryde, NSW, Australia, serial number: 820121). According to ISO 10993-12:2021 [36], an extract of the test item (Figure 11) in culture medium was prepared by immersing the device into the medium. After extraction, the extracts of the controls and the test item were clear, with no observable particulates. The extracts were not filtered, centrifuged, or otherwise altered and were used immediately after preparation. Triplicate monolayers of L929 cells were dosed with 600 µL of extracts and incubated in the presence of 5 ± 0.1% CO₂ for 24 to 26 h. A staining solution was prepared, just before use, by mixing Trypan Blue solution with supplemented MEM in a 1:1 ratio. Following the incubation, 100 μL of the prepared staining solution was dispensed into each well. Afterwards, the cytotoxicity was assessed via microscopic observations.

3. Results

3.1. Tensile Test

Eight sets of samples were tested, examples of which are shown in the photo below (Figure 12).

Based on the static tensile test, charts were created to determine characteristic points on the stress–strain curve, thanks to which it was possible to describe the mechanical properties of a given material. These points are the elastic limit, the yield point, and the tensile strength. The obtained waveforms for each set are presented in the charts below (Figure 13).

Based on the results obtained (

Table 4. Mechanical properties obtained from quasi-static tensile testing.

Material		Elastic Limit		Yield Strength		Ultimate Tensile Strength
		Stress [MPa]	Strain [mm/mm]	Stress [MPa]	Strain [mm/mm]	Stress [MPa]	Strain [mm/mm]
1.	PETG Carbon	20.54 ± 0.48	0.008 ± 0.04	29.18 ± 0.38	0.01 ± 0.01	60.09 ± 1.80	0.063 ± 0.00
2.	FiberFlex 40D	1.31 ± 0.40	0.050 ± 0.01	2.41 ± 0.03	0.22 ± 0.01	8.71 ± 0.55	5.090 ± 0.30
3.	PETG Sunlu	13.18 ± 0.17	0.011 ± 0.01	15.08 ± 0.18	0.01 ± 0.01	31.49 ± 1.07	0.029 ± 0.01
4.	HIPS Devil Design	9.44 ± 0.03	0.014 ± 0.01	12.73 ± 0.32	0.02 ± 0.01	10.83 ± 0.17	0.366 ± 0.01
5.	TPU	0.69 ± 0.15	0.043 ± 0.01	1.97 ± 0.41	0.28 ± 0.02	10.96 ± 1.22	2.960 ± 0.42
6.	PA+15CF	14.71 ± 0.38	0.010 ± 0.01	20.03 ± 0.38	0.01 ± 0.01	56.57 ± 1.09	0.073 ± 0.01
7.	PETG 3DActive	18.01 ± 0.18	0.018 ± 0.01	22.01 ± 0.24	0.02 ± 0.01	43.38 ± 0.68	0.061 ± 0.01
8.	PLA 3DActive	21.95 ± 0.17	0.014 ± 0.01	48.89 ± 0.89	0.03 ± 0.01	43.89 ± 1.16	0.110 ± 0.02

), it can be concluded that the highest stress value (60.09 MPa) was obtained for the PETG Carbon material, which exhibits elastic–plastic behavior without a clear yield point. PA+15CF, PETG 3DActive, HIPS, and PLA also exhibit similar behavior, which is typical of polymer materials. Unlike its two counterparts, PETG Sunlu does not exhibit a ductile range, exhibiting brittle behavior, cracking when a stress of 31.49 MPa is reached. An interesting case is the behavior of the TPU and FiberFlex 40D materials, which exhibit elastomeric behavior by deforming to a significant extent compared to the other materials. FiberFlex 40D (Fiberlogy, Brzezie, Poland) obtained the highest strain value of 509%, at which the sample broke, and a stress of 8.71 MPa was recorded. Important in the context of materials engineering, these recorded values allow for their initial assessment in relation to potential applications in industry, rapid prototyping, and further development in the context of materials engineering.

3.2. SEM Scrap Observation After Tensile Test

Figure 14, Figure 15, Figure 16 and Figure 17 show the surface topography after a static tensile test of samples made of PLA, PET-G, and HIPS without additional strengthening in the form of fillers. Figure 14 shows a view of the topography of the fracture surface of a sample made of the PLA 3DActive filament.

Figure 14 shows a system of typical faults visible in the surface topography related to the passage of the crack front through an area of locally developed deformations. On a micro-scale, systems of low faults in the fractures of the layered filament are clearly visible throughout the entire surface. On a macro-scale, in the area with the “fill” printing function, visible material extrusions are traces of plastic deformations. Figure 15 shows a view of the topography of the fracture surface of a sample made of the PET-G Sunlu filament.

Figure 15 shows the crack focus at the edge of the sample, and an extensive system of radially arranged high faults associated with the passage of the crack front are visible in the surface topography.

Figure 16 shows a view of the topography of the fracture surface of a sample made of the PET-G 3DActive filament.

The sample visible in Figure 16 is set at an angle in the microscope chamber. A system of faults visible in the surface topography is related to the passage of the crack front from the edge of the sample through the spatially expanded remaining deformation area. On the surface near the upper edge, systems of low faults in the breaks of the layered filament are clearly visible. In the central areas of the sample, local wedge-like deformations with a plastically extruded apex are visible. In the wall area, there are numerous empty spaces between the applied filament layers, with fragments without visible fusions with the adjacent material paths.

Figure 17 shows a view of the topography of the fracture surface of a sample made of the HIPS filament by Devil Design.

Figure 17 shows a system of typical high faults visible in the surface topography that are related to the passage of the crack front through the central area of the sample, with local deformations of individual HIPS filament paths. A visible change in the direction of the crack movement in the layers deposited within the wall can been clearly seen (red arrow). Also, there are clearly visible patterns of low faults on the entire surface in areas with well-grafted filament paths during printing, especially in wall areas where there are some empty spaces between the applied filament layers. The fracture visible in Figure 17 shows the surface topography with the most distinct structure showing a change in the direction of cracking during the tensile test.

A common feature of the resulting scrap (Figure 14, Figure 15, Figure 16 and Figure 17) was the presence, on a macro-scale, of high and low faults constituting a trace of the fracture passage, which created rivers and basins spreading from the edges of the samples through their central part to the opposite edges. At the same time, at the magnification used (40×), the system of low faults was clearly visible, also within a few or a dozen well-connected individual paths of the printed filament. In the case of the PET-G filament from 3DActive (Figure 16), unlike the others, in the central area of the scrap, the broken connectors of the plastically deformed phase are very clearly visible, creating forms that locally resemble wedges or “shark teeth”. The observations made in Figure 14, Figure 15 and Figure 17 lead to defining the topography of the surfaces formed after the static tensile test as ductile–brittle fractures. In the case of the PET-G material from Sunlu (Figure 17), the surface topography in the macroscopic observations resembles the structure of a low-energy brittle fracture without clearly visible traces of plastic deformation of the material.

The next group of materials subjected to the static tensile test were the carbon fiber composites obtained after FDM printing from PET-G Carbon and PA+15CF filaments. Figure 18 shows a view of the topography of the fracture surface of a sample made of the PET-G Carbon filament.

As can be seen in Figure 19, there are numerous material stretches visible in the surface topography with broken (cracked) carbon fibers of the filler stuck in the matrix. This is an example of a fracture structure without a clearly defined fracture front.

Figure 20 shows a view of the topography of the fracture surface of a sample made of the PA+15CF filament.

Similar to PET-G Carbon (Figure 19), scrap observations, shown in Figure 20, of a PA+15 CF sample fracture structure are visible in the surface topography without a clearly defined crack front. On the surface, there are numerous oriented CF carbon fibers residing in the matrix (15%).

The conclusions from the observations of fractures presented in Figure 19 and Figure 20 are consistent with the results of tensile strength tests in a static tensile test (Table 4). Compared to other materials, the highest R_m strength values were obtained in composite samples containing carbon fiber reinforcement in the matrix. The performed tests for the mechanical parameters (tension and bending) highlight the validity of their use despite the very difficult conditions of producing elements (samples) from these materials using the FDM method. SEM observations (Figure 19 and Figure 20) in the topography module did not clearly reveal an area in the surface structure where the beginning of the sample cracking process or the directions of crack propagation in the cross-section could be identified. Differences in the heights of faults are visible between the areas constituting the outer walls and the topography of the fracture in the central part of the samples produced with the infill printing function.

Below, Figure 18 shows SEM views without the topography observation function to reveal details such as the arrangement of filler fibers and areas of possible deformations.

In both Figure 18a,b, the details of the fracture surface within the printed walls have the character of cavities in which the deformation was initiated and, thus, resemble voids usually occurring around hard dispersion particles (fibers). Most likely, the occurrence of voids in the surface topology is influenced by the arrangement of the fibers, which, in this area, lie parallel to the direction of operation of the printing nozzle, i.e., perpendicular to the observed surface. In the case of the central area of the samples, where in both materials the direction of the printing nozzle changed in each subsequent layer, there are only clearly visible faults and fragments and traces of fibers “lying” parallel to the observed surface. It is worth noting that the fibers protruding from the PA surface (Figure 18b and Figure 20) did not break evenly with the matrix material, and their being stuck above the scrap surface may indicate high strength exceeding the adhesion forces occurring between them and the PA matrix and the matrix itself. The process of pulling fibers from PA was accompanied by plastic deformations, thus creating an extensive surface topography. Observations of the fracture surface structures of both tested composite materials with fibers residing in them to a greater or lesser extent explain the slightly different course of the tensile test for each of them, which may be related to many factors (share and structure of fibers, matrix material, and printing parameters). The tensile strength value (Table 4) of the PA+15CF filament composite was slightly reduced (R_m = 56.57 MPa) compared to the PETG Carbon material, for which the tensile strength was 60.09 MPa; however, the observed topography of both scraps with low and high faults and details such as voids and material extrusions is similar and resembles ductile–brittle fractures.

3.3. Three-Point Bending

For the three-point bending test, the sample geometry conformed to the PN-EN ISO 178 standard [37] was used. The same material set was tested as in the static tensile test. Example samples after the test are presented in Figure 21.

Testing the materials using the three-point bending method allowed for the creation of stress–displacement graphs, on the basis of which basic information was obtained about the mechanical properties of materials in the context of bending strength. These properties were determined on the basis of the points read on the charts, which included the bending strength and maximum elongation. The obtained curves are presented in the charts below (Figure 22).

Based on the read values (

Table 5. Mechanical properties obtained from three-point bending test.

Material		Ry Yield Stress [MPa]	Elongation at Yield [mm]	Ru Ultimate Tensile Stress [MPa]	Elongation at Break [mm]
1.	PETG Carbon	113.69 ± 0.52	2.14 ± 0.004	112.59 ± 0.79	2.38 ± 0.001
2.	FiberFlex 40D	-	-	-	-
3.	PETG Sunlu	36.55 ± 0.31	1.66 ± 0.020	2.41 ± 0.03	0.22 ± 0.005
4.	HIPS Devil Design	22.96 ± 0.73	2.33 ± 0.020	9.14 ± 0.02	5.31 ± 0.010
5.	TPU	-	-	-	-
6.	PA+15CF	98.26 ± 0.20	3.55 ± 0.040	1.99 ± 0.03	11.86 ± 0.080
7.	PETG 3DActive	74.02 ± 0.08	3.12 ±0.020	70.77 ± 0.08	3.35 ± 0.060
8.	PLA 3DActive	66.02 ± 0.10	1.56 ± 0.050	1.20 ± 0.10	8.48 ± 0.150

), PETG Carbon had the highest bending strength value (113.69 MPa), but it had the second lowest maximum elongation compared to other tested materials. Materials such as PA+15CF, PLA, and HIPS exhibited behavior similar to that of PETG Carbon. An exception to this were samples made of PETG Sunlu and PETG 3DActive, which cracked brittlely after reaching the maximum stress. An interesting case was the tested materials TPU and FiberFlex 40D, which were not destroyed after reaching the maximum displacement of 50 mm for the research equipment used. After removing the stress, these samples partially returned to their original shape, suggesting the elastomeric nature of these materials.

3.4. SEM Scrap Observation After Three-Point Bending Test

SEM observations of the surface conditions of fractures obtained after strength tests were carried out at 40× and 250× magnifications on samples coated (sprayed) with graphite. Reference areas located near the center of the obtained fractures were selected for testing. The SEM topography of the fracture surfaces obtained after the static three-point bending test for the selected samples is shown in Figure 23. The drawing also shows the observation areas selected for research.

On each of the observed sample fracture surfaces, the influence of the direction of the print paths on the structure and surface topography of the obtained fractures was revealed. The band structure of the resulting fractures is best revealed by observations carried out at low magnifications of 40× to 60×. Therefore, in the case of fracture surfaces of samples without the presence of fibers in the matrix, such as PETG Sunlu, PETG 3DActive, HIPS Devil Design, and PLA, a system of faults related to the passage of the crack front through areas of local deformation is visible on the fracture surfaces of the samples. They are formed along the (vertical) print paths. An example of a fault pattern along the (vertically arranged) print paths of the HIPS Devil Design sample is shown in Figure 24a. A similar, although less transparent, band arrangement of faults is characterized by breakthroughs in samples containing the fibers PETG Carbon and PA+15CF Rosa 3D. An example of the fracture surface of the PETG Carbon sample in a general view obtained at 40× magnification is shown in Figure 24b.

Faults forming extensive planes are observed on the scrap surfaces of the PETG Sunlu samples compared to the surfaces of the PETG 3DActive samples, where they are clearly denser (Figure 25). In both samples, the transition between adjacent planes is separated by a sharply outlined edge. Moreover, numerous micro-faults are visible on the fault surfaces, creating a characteristic striation effect, and their number clearly increases at the fractures of the PETG 3DActive samples. In both samples, the fracture characteristics are appropriate for materials characterized by brittle fractures, which corresponds to the results obtained in the σ-ε diagram. It should also be noted that the increased strength of the PETG 3DActive sample over PETG Sunlu is consistent with the increased number of dislocations observed on its surface.

On the entire fracture surface of the PLA 3DActive samples, clear material pulls are observed (marked in the drawing) in the direction of the tearing force, occurring mainly at the border of the print paths (Figure 26b). These extrusions in their structure correspond to materials with increased ductility. In turn, on the surface of the HIPS Devil Design samples (Figure 26a), in the marked areas outside the faulted bands, the separation surfaces after detaching the print path are shown. In this case, the high susceptibility to detachment of the print paths had a negative impact on the strength properties.

The highest values of bending strength were found in composite samples containing fiber reinforcement in the matrix. High values of bending strength Rg exceeding 113 MPa are reflected in the structure of the fracture surfaces of the PETG Carbon samples, exemplified in Figure 27a. SEM observations of the surface revealed numerous material pull-outs in the structure of the fracture surface with broken (cracked) filler fibers stuck in the matrix (indicated by arrows). This nature of the composite fracture confirms the strong bond at the interface of the PETG matrix material with carbon fiber reinforcement. The strong adhesive bonding of the components contributed to the increase in the strength Rg while limiting the ability to deform. A similar nature of destruction is observed on the failure surfaces of samples made of PA+15CF Rosa 3D material. In this case, however, the scrap formation process takes place with numerous pull-outs of CF fibers from the PA matrix material. In Figure 27b, the arrows indicate places in the matrix after pulling out the fibers. Therefore, it can be concluded that in relation to the PETG Carbon material, the PA+15CF material is characterized by lower adhesion and a weakened connection at the interface of the CF fibers with the PA matrix. Observations of the structure of the fracture surface explain, in this case, the reduced value of bending strength Rg compared to the PETG Carbon material.

Observations of the fractures of samples made of filaments highlighted the influence of the quality of workmanship in the 3D printing of samples on the fractures created, in which it is possible to distinguish the layers of material applied to create the walls and the layers forming the internal filling of the cross-section. The variety of these structures (directionality) and the materials used for the filaments leads to spatially extensive destruction. The observed local discontinuity of successively applied layers may result in worse-than-expected mechanical parameters than in materials with a more uniform structure.

3.5. Numerical Simulations of Tensile Test

The determination of Johnson-Cook (J-C) parameters through experimental testing and numerical simulation validation is a well-established methodology in materials research. It has been successfully applied across various material classes, ensuring that the mechanical behavior captured by the model is consistent with experimental observations. Studies such as those by Chen et al. for Ti–6Al–4V [42], Dorogoy and Rittel for Ti6Al4V [43], and Jang et al. using SHPB tests [44] demonstrate the robustness of this approach for metals and alloys. Similar methodologies have also been extended to thermoplastics, as shown in the work of Amani et al. for thermoplastic composites [45] and Dehgolan et al. for structural steel [46].

However, while these methods are extensively applied to metals and conventionally processed materials, research addressing the determination of J-C parameters specifically for 3D-printed materials, such as PETG, PLA, PA+15CF, or HIPS, remains sparse. This gap highlights the novelty of the current study, which applies tensile and flexural testing combined with numerical simulations to bridge this knowledge gap for additively manufactured thermoplastics.

Below (Figure 28) are graphs comparing the experimental results from the strength tests of filament samples with the results of simulation tests in the Abaqus/Explicit computing environment. The results of the tensile test were compared after the application of previously developed parameters in the Johnson–Cook material model.

Experimental studies and numerical simulations made it possible to assess the suitability of the Johnson–Cook model for future numerical analyses of components made from commercially available filaments, and indicated the possible parameters needed to use the model. In order to verify the compatibility of the curves obtained from numerical simulations and experimental studies, a method was used to compare the areas bounded by these curves in a designated area. The boundaries of the areas were determined at the points of intersection of the curves. Then, the obtained values of area differences for each area were summed, without considering the sign, and related to the area under the curve obtained experimentally. The results can be seen below (Figure 29).

The results show a high degree of agreement between the curves obtained via numerical analysis and those determined via experimental testing for materials such as FiberFlex 40D, PETG, HIPS, TPU, and PA+15CF Rosa 3D. Figure 29 illustrates that the maximum fitting error for the mentioned materials does not exceed the value of 2.6%. This state of affairs is due to the monotonic nature of the tested materials. In the case of each sample from the listed materials, the phenomenon of plastic strengthening was observed, which resulted in a powerfully increasing dependence of stress on strain. This behavior predisposes the use of the Johnson–Cook model due to the first term of the equation appearing in this model, which is also of a power-law nature. Minor deviations in the fit for the TPU filament in the yield region result from the difficulty of representing a smooth transition from elastic to plastic range. Therefore, an abrupt change in the angle of the curve from the numerical analyses is apparent in that area. For the HIPS material, it was impossible to map the characteristic willow, after which the polymer in the strain range of 0.025 to 0.070 underwent plastic softening and then, until rupture, manifested a perfectly plastic character. Therefore, the material parameters related to the exponent of strengthening and the coefficient of strengthening were established as zero values. The mapping of the PET-G Carbon material also showed an acceptable level of error. However, it was noticeably higher than for the previously discussed filaments. The reason for this is, as in the TPU material, the smooth transition from the elastic to the plastic state. The difference is that the transition area is wider in relation to the total deformation of the material than in the case of the TPU filament, resulting in an increased value of the mapping error.

The conducted analyses showed that the Johnson–Cook model is not adequate to describe the behavior of the PLA 3D and PET-G materials. The lack of a monotonic character as well as the occurrence of plastic shrinkage effectively exclude the further use of the adopted material model for their description in numerical analyses. This is due to the too-complex nature of the cited filaments, which the formulations of the given model are not able to describe with a satisfactory level of error. Despite the indicated error of 6.84% for PET-G and 4.70% for PLA 3, the study showed the completely incompatible nature of the curves obtained numerically with the experimental curves.

3.6. Numerical Simulations of Three-Point Bending

In the case of the three-point bending test, the values obtained from numerical simulations do not coincide with the experimental results. This is due to the fact that in the Johnson–Cook model, the stress and strain values are obtained during tension, which causes the sample to be subjected to a more complex bending load condition, which results in different stress and displacement values, which lead to the sample failing. This is also influenced by the geometric notch (cutting point) of the sample. Additionally, it is also worth mentioning that the system is much less stable. In the case of simulation, the system suddenly changes its form; that is, some finite elements are destroyed, so the measurement of stresses in these elements is suspended. This also introduces significant differences in the way the material is destroyed. Fragments of the sample near the geometric notch are destroyed much faster than elements in the remaining volume of the sample, and the differences and fluctuations in the values obtained as a result of numerical simulations for the selected areas of the sample are shown in the example of the PLA material (Figure 30) and two places: next to the geometric notch and close to contact with the sample loading indenter.

3.7. Cytotoxicity Tests

The potential of cytotoxicity of the evaluated samples (

Table 6. Obtained filament material parameters based on experimental research.

Sample	PETF Carbon	FiberFlex 40D	PETG	HIPS	TPU	PA+15CF Rosa	PETG 3DActive	PLA
Blank	0	0	0	0	0	0	0	0
Negative control	0	0	0	0	0	0	0	0
Positive control	4	4	4	4	4	4	4	4
Test item	Non-cytotoxic	Cytotoxic	Non-cytotoxic	Non-cytotoxic	Non-cytotoxic	Non-cytotoxic	Non-cytotoxic	Non-cytotoxic

) was investigated using a qualitative, in vitro cytotoxicity MEM elution assay in accordance with ISO 10993-5:2009 [28] and ISO 10993-12:2021 [36]. Following the suggestions of ISO 10993-5 Clause 4.1, the test was performed on an extract of the test sample. Sterile complete culture medium was chosen because this extraction vehicle is recommended by ISO 10993-5 Clause 4.2.2, as it combines both polar (basic culture medium) and non-polar (10% fetal bovine serum) extraction conditions, allowing for the investigation of all soluble hydrophilic and lipophilic substances from the material of the product as well as any surface contaminations present on the final product.

The cytotoxic potential was assessed based on PN-EN ISO 10993-5:2009, Biological evaluation of medical devices, Part 5: Tests for in vitro cytotoxicity, Table 1: “Qualitative morphological grading of cytotoxicity of extracts”. The achievement of a numerical grade of 3 or greater is considered a cytotoxic effect. The acceptance criteria were as follows: blank grade < 3 positive control grade ≥ 3 negative control grade < 3.

On the basis of the results interpreted according to PN-EN ISO 10993-5:2009, the test item Sample Nos. 1 and 3–8 (PETG Carbon, PETG, HIPS, TPU, PA+15CF, PETG, and PLA) should be considered non-cytotoxic. On the basis of the results interpreted according to PN-EN ISO 10993-5:2009, the test item Sample No. 2, FiberFlex 40D Fiberlogy, should be considered cytotoxic.

4. Discussion

Based on the mechanical tests and numerical simulations, several important conclusions can be drawn regarding the material properties of filaments used in FDM technology. The basic tensile test of the materials allowed for the determination of key mechanical properties, which enabled the initial development of a rheological model of the materials. These results are crucial for further modeling and predicting the behaviors of materials during mechanical processing.

Thanks to the research, printing parameters, such as temperatures, have been optimized in relation to the recommendations of filament manufacturers. It turned out that these recommendations were often inaccurate, which could affect the quality of prints and the mechanical properties of the final products.

Carrying out a tensile test made it possible to determine the parameters of the Johnson–Cook model and determine the limit strain at which the structure becomes discontinuous. Thanks to this, the fracture mechanics parameter was also identified, which is important for assessing the durability of materials in engineering applications. The experimental tensile strength value of the standard material, PLA, was close to the values achieved in other studies [25,47], which suggests an optimal outcome for the other tested materials.

Numerical simulations showed high convergence with real results, reaching 97%, which proves that the models fit reality very well. The largest differences, exceeding 3%, were observed in the case of materials such as PETG Carbon, PET-G, and PLA. However, the Johnson–Cook model did not perform well in terms of simulating three-point bending. The model based on a uniaxial tensile test is not suitable for tests involving more complex material stress states, and therefore, it is not recommended to use this model in this case.

The analysis of the fracture microstructure using an SEM allowed for the observation of cracks on individual material layers, which results directly from the characteristics of the FDM printing technology. Such a phenomenon was observed in other studies [48] and is key to understanding the cracking mechanisms of these materials and optimizing the printing process to minimize structural defects.

Based upon the above results and evaluation arguments, it is concluded that the investigational samples PETG Carbon, PETG, HIPS, TPU, PA+15CF, PETG, and PLA have no cytotoxic properties in terms of ISO 10993-5, relating to potentially toxic substances being released from the product under investigation, which is correlated with the outcomes for PLA, PETG, and PA in other studies [49,50]. These results suggest that the materials described can be used in applications requiring direct contact with living tissue. At the same time, it was found that the sample FiberFlex 40D was found to be toxic and should not be used as a biomaterial.

5. Conclusions

The experimental and numerical studies demonstrated that the Johnson–Cook material model effectively predicts the mechanical behaviors of certain commercially available filaments used in additive manufacturing. By comparing the stress–strain curves from simulations and experiments, we quantitatively assessed the model’s accuracy across different materials. A strong agreement was found for monotonic materials like FiberFlex 40D, PETG, HIPS, TPU, and PA+15CF Rosa 3D, with maximum fitting errors not exceeding 2.6%. This correlation is due to their monotonic stress–strain relationships, which the Johnson–Cook model captures well through its power-law formulation. The parameters determined in this study can be utilized in simulations to enhance predictive capabilities. Minor discrepancies in materials like TPU and PET-G Carbon arose from challenges in modeling the smooth transition from elastic to plastic behavior, but the overall errors remained acceptable. However, the model was inadequate for non-monotonic materials like PLA 3D and PET-G, resulting in higher errors and incompatible numerical and experimental curves.

Cytotoxicity tests, which were conducted under ISO 10993-5 and 10993-12 standards, confirmed that materials like PETG Carbon, PETG, HIPS, TPU, PA+15CF, and PLA are non-cytotoxic and potentially safe for biomedical applications. In contrast, FiberFlex 40D exhibited cytotoxic properties, making it unsuitable for contact with biological tissues without modification.

Scanning electron microscopy provided insights into the fracture mechanisms of various filament materials used in additive manufacturing. The material composition significantly influenced fracture behavior. Unreinforced materials like PLA 3DActive, PET-G Sunlu, PET-G 3DActive, and HIPS Devil Design exhibited fracture surfaces with fault systems indicative of combined ductile and brittle mechanisms, forming patterns resembling rivers and basins. This suggests that both brittle and ductile failure modes contribute to their mechanical performance. In contrast, fiber-reinforced composites such as PET-G Carbon and PA+15CF displayed distinct fracture characteristics. PET-G Carbon samples showed broken carbon fibers embedded in the matrix, indicating strong fiber–matrix adhesion and effective load transfer, enhancing the mechanical strength and stiffness. PA+15CF samples exhibited fiber pull-out from the matrix, pointing to weaker fiber–matrix adhesion; while this absorbs energy during fracture, it may reduce overall strength compared to composites with stronger interfacial bonding. The quality of 3D printing significantly affected fracture behavior. Well-bonded filament paths improved the mechanical properties due to efficient stress distribution and resistance to crack initiation, whereas poor layer adhesion and voids acted as stress concentrators, reducing the strength and facilitating crack growth along weak interfaces. The print path directionality influenced crack propagation patterns. In unreinforced materials, cracks followed stepped patterns along poorly bonded layers, reflecting a combination of brittle fractures at weak interfaces and ductile deformation in well-bonded regions. For fiber-reinforced composites, crack propagation was more linear, with embedded fibers bridging cracks and reinforcing the matrix, contributing to higher strength and stiffness. The SEM observations correlated with the mechanical test results. Materials exhibiting ductile fracture features, such as material pull-outs and plastic deformation, demonstrated higher elongation and toughness, making them suitable for applications requiring energy absorption and flexibility. Fiber-reinforced materials achieved higher tensile and bending strengths, consistent with the effective fiber reinforcement and strong fiber–matrix adhesion observed in fracture surfaces.

Future studies should focus on accelerated aging tests and environmental simulations to better understand the degradation pathways and mechanical fatigue of these materials over time. Additionally, optimizing numerical models to capture complex stress behaviors will further enhance the predictive accuracy and reliability of these materials in real-world applications.