Experimental and numerical analysis of the effect of strut orientation on the bending and compression behavior of FBCCZ lattice structure

ABSTRACT

In this study, the mechanical behavior and energy absorption performance of FBCCZ and FBCCY lattice structures with identical relative density were investigated experimentally and numerically under compression and three-point bending loading conditions. The lattice specimens were fabricated using stereolithography and subjected to compression and bending experiments according to ASTM standards. Finite element analyses were performed using ANSYS Static Structural and LS-DYNA solvers, and a mesh convergence study was conducted to verify the reliability of the numerical model. The experimental and numerical results demonstrated that strut orientation significantly affects deformation behavior, load-carrying capacity, and energy absorption capability of the lattice structures. Under compressive loading, FBCCZ-oriented specimens exhibited approximately 30.9% higher experimental energy absorption and 33.4% higher numerical energy absorption compared to FBCCY-oriented specimens. In contrast, FBCCY-oriented structures demonstrated superior load-carrying and energy absorption behavior under three-point bending conditions. Different collapse mechanisms were also observed depending on strut orientation; collapse initiated from the top region in FBCCY structures, whereas deformation began from the bottom region in FBCCZ structures. The numerical analyses showed good agreement with the experimental results under both loading conditions. The results indicate that strut orientation should be considered a crucial design parameter for lightweight lattice structures under different loading conditions.

Keywords:
Lattice structures; Energy absorption; Stereolithography; Finite element analysis.

1. INTRODUCTION

Lattice structures that are lightweight and offer high specific strength have been widely studied for energy absorption in the automotive and aerospace fields, owing to their benefits, including improved fuel consumption, reduced pollution, and increased crashworthiness [¹, ², ³, ⁴, ⁵].

Lattice structures can be categorised by their mechanical properties; in the literature, these structures typically fall into two categories: stretch-dominated and bending-dominated. Stretch-dominated structures are those whose deformation is dependent on axial load carrying mechanisms, whereas in the case of bending-dominated structures, the deformation depends on bending [⁶]. To differentiate between these structures, classify these structures, and analyze their mechanical response, a common method is the Maxwell criterion (M = b – 3j + 6), where M is the value calculated depending upon the number of struts and joints present in the structure [⁷, ⁸]. However, many researchers have shown that the Maxwell criterion does not necessarily provide a clear basis for determining the deformation mechanism of lattice structures and that this theory-based approach may be inadequate in certain cases, especially when complex geometry is involved [⁹]. The mechanical properties of lattice structures are influenced by several parameters, including relative density, unit cell geometry, strut shape and size, cell arrangement, material properties, and manufacturing process [¹⁰, ¹¹, ¹²]. This indicates that variations in cell configuration play a decisive role in determining load-bearing capacity and deformation behavior. In strut-based lattice structures, increasing the number of unit cells and incorporating additional supporting elements enhance compressive strength and energy absorption capacity [¹³].

Stereolithography (SLA) is a vat photopolymerization technique based on the layer-by-layer curing of a liquid photosensitive resin through vector scanning by a laser beam. This method enables the fabrication of three-dimensional polymer structures with high dimensional accuracy and fine detail resolution [¹⁴, ¹⁵].

Advances in additive manufacturing technologies have made it possible to produce complex cellular geometries. Vat photopolymerization techniques, including SLA, digital light processing (DLP), and masked stereolithography (mSLA), are widely used in the fabrication of lattice structures due to their high resolution and surface quality [¹⁶]. However, it has been reported in the literature that printing orientation and layer thickness significantly affect mechanical properties [¹⁷]. In addition, washing duration and post-curing parameters have been shown to influence tensile and flexural strength [¹⁸], while curing time and temperature alter the overall mechanical performance [¹⁹]. Therefore, manufacturing and post-processing parameters must be considered together when evaluating the mechanical performance of lattice structures.

In the literature, the energy absorption performance of various unit cell topologies such as BCC, octet, pyramidal, and triply periodic minimal surface (TPMS)-based structures has been extensively compared [², ³, ¹³]. However, comparative studies focusing on structures with the same fundamental cell morphology but differing only in axial reinforcement orientation remain limited.

In the literature, the mechanical response of cellular structures is evaluated in terms of parameters such as unit cell geometry, strut thickness, and configuration [²⁰, ²¹]. In this context, it has been experimentally demonstrated that vertical Z-strut reinforcement aligned with the loading direction alters the deformation mode of FBCC unit cells and significantly influences their energy absorption behavior [²²].

Recent studies have increasingly focused on bio-inspired structural designs to improve crashworthiness and energy absorption performance. Various bio-inspired and hybrid configurations have demonstrated that geometric design and reinforcement arrangement can significantly influence deformation mechanisms and energy absorption under different loading conditions [²³, ²⁴, ²⁵]. Such studies demonstrate the growing interest in geometry-based approaches for the development of lightweight energy-absorbing structures.

The contributions of this research to the existing literature include the experimental and numerical analysis of the impact of strut direction on the energy absorption performance of FBCCZ and FBCCY configurations, both with the same relative density and unit cell structure, subjected to compressive and bending loading. The results obtained from the study can be valuable to designers in the automotive industry for the manufacture of lightweight cellular cores.

2. MATERIALS AND METHODS

2.1. Specimen design and geometrical properties

The lattice model under consideration was designed as an FBCCZ type, as shown in Figure 1. The structure was modeled in SolidWorks 2021. The length of the unit cell edge was 8 mm, chosen according to the geometrical parameters used in previous studies [²⁶, ²⁷, ²⁸]. A strut diameter of 1.5 mm was used to fabricate the lattice structure. Compression test specimens were constructed by repeating unit cells in a 3 × 3 × 6 spatial arrangement, comprising a total of 54 unit cells. Dimensional parameters for compression specimens were 24 mm × 24 mm cross-sectional area with a height of 48 mm, giving a ratio of height to width of 2:1 according to ASTM D695 standard [²⁹].

Figure 1
(a) FBCCZ unit cell structure, (b) Z-oriented structure, and (c) Y-oriented structure.

In the literature, the number of struts and nodes for the FBCCZ topology are reported as b = 36 and j = 15, respectively, and the Maxwell number is calculated as M = −3 according to the relation M = b − 3j + 6. This value indicates that the structure exhibits a bending-dominated behavior [⁹].

In order to evaluate the effect of loading direction on mechanical behavior, the specimens were prepared in two different orientations while keeping the unit cell geometry and strut dimensions constant. In the FBCCZ configuration, loading was applied along the Z-axis, whereas the FBCCY specimens were generated by changing only the replication direction to align with the Y-axis. The unit cell geometry remained unchanged. Therefore, the relative density of FBCCZ and FBCCY specimens is equivalent by design. In the following sections, these models are designated as Type 1 (FBCCZ) and Type 2 (FBCCY) for ease of comparison.

The measured mass of the specimen was determined to be approximately 11.5% greater (11.955 g) than that of the computer-aided design (CAD) model’s prediction of 10.58 g. The difference in mass can be explained by the presence of additional resin retained inside voids in the model. The relative density was calculated as the ratio of the lattice structure mass obtained from the CAD model to the mass of a solid block with the same external dimensions (31.8 g), resulting in a relative density of 33.27%.

2.2. Manufacturing process

The specimens were fabricated using Anycubic Tough Resin 2.0 photopolymer resin. According to the manufacturer’s technical datasheet, the density of the resin is reported to be in the range of 1.10–1.15 g/cm³. The models designed in SolidWorks 2021 were exported in .stl format and processed in the printer’s slicing software, Photon Workshop. Support structures were generated during the slicing process. The fabrication process was carried out using an Anycubic Photon Mono M5s Pro printer, and the printing time for each specimen was recorded as approximately 3 hours and 45 minutes.

The specimens were fabricated using the printing parameters recommended by the manufacturer. The applied printing parameters are presented in Table 1.

Table 1
Manufacturer-recommended SLA printing parameters for Tough Resin 2.0 (Photon Mono M5s Pro).

PARAMETER	VALUE
Layer Thickness	0.05 mm
Normal exposure time	3.600 s
Bottom layers	5
Bottom exposure time	30 s
Off time	1 s
Bottom off time	0.5 s
Z lift distance	8.000 mm
Z lift speed	6 mm/s
Z retract speed	6 mm/s

During the manufacturing process, all specimens were prepared under identical environmental conditions. The printing temperature was maintained within the range of 25–30°C, and the printing parameters were kept constant to ensure experimental comparability. No secondary (UV) post-curing process was applied after printing. The specimens were stored under ambient conditions and subjected to mechanical testing within 15 days after fabrication. Accordingly, the specimens were evaluated in the as-printed condition, with identical processing and testing conditions maintained for both lattice orientations to ensure a consistent comparative assessment. The printing preparation stage and fabricated specimens are shown in Figure 2.

Figure 2
(a) Printing preparation stage and (b) fabricated specimens of FBCCZ and FBCCY structures.

2.3. Experimental setup and mechanical testing

In order to determine the mechanical properties of the resin in the as-printed condition, tensile specimens were modeled in SolidWorks 2021 (Figure 3) and fabricated in accordance with the ASTM D638 Type I standard for tensile testing of plastic materials [³⁰]. Tensile tests were conducted at a constant crosshead speed using a Shimadzu AGS-J 10 kN universal testing machine, and the data were recorded simultaneously. Based on the obtained stress–strain curves, the Young’s modulus, yield stress, and tangent modulus were determined as 500 MPa, 15 MPa, and 15 MPa, respectively, and subsequently used to define a Bilinear Isotropic Hardening material model in the finite element analyses.

Figure 3
(a) Tensile specimens prepared according to ASTM D638 Type I standard and (b) specimen geometry created in SolidWorks.

For each loading condition, three independent specimens of each lattice orientation (Type 1 and Type 2) were tested, and the experimental energy absorption results are reported as mean ± standard deviation. Compression tests were conducted in accordance with the ASTM D695 standard using a Shimadzu AGS-J 10 kN universal testing machine at a constant loading rate of 1 mm/min. During the tests, force–displacement data were continuously recorded. The specimens were subjected to compression until the desired displacement was attained, and the tests were stopped at about 50% strain.

Three-point bending tests were performed according to the ASTM D790 standard protocol. The specimens had dimensions 24 × 24 × 240 mm, while the span between the supports was 192 mm (refer to Figure 4). The span length of the specimens was chosen to have an overhang of at least 10% past the support span lengths on either side, which fulfills the standards set by the ASTM standard protocol [³¹]. A loading rate of 2 mm/min was chosen, and the force-displacement relationship was measured during the tests. As no fractures occurred in the specimens, the tests were stopped at a displacement of 10 mm.

Figure 4
Three-point bending test setup: (a) Type 1 specimen and (b) Type 2 specimen.

2.4. Finite element analysis

The numerical simulation for the three-point bending test was performed in ANSYS Workbench 2021 R1 using the Static Structural tool, consistent with a previous numerical study in which Static Structural analysis was used for three-point bending, while explicit analysis was adopted for compression [³²]. Furthermore, comparative analyses of three-point bending have shown that both approaches can be used, while the implicit solution can require substantially lower computational time than the explicit solution under certain loading conditions [³³]. Therefore, Static Structural was preferred for the three-point bending analysis in the present study. An appropriate model for the experiment was created, and the lower rollers supporting the beam were treated as fixed supports, with all DOF constrained. To achieve a total displacement of 10 mm within the time frame of 0–300 s, the displacement was specified as Tabular Data for the upper roller. The diameters of the loading and support rollers were set to 25 mm.

To better simulate the regions where stresses concentrate under load, the Sphere of Influence technique was used, with a grid spacing of 1.5 mm over a 30 mm radius. Force–displacement curves were generated using the reaction force values obtained from the upper loading roller, and the results were compared with the experimental data. The solver type used in the analyses, along with the number of nodes and elements in the model, are provided in Table 2.

Table 2
Number of nodes and elements for Type 1 and Type 2 lattice structures.

SPECIMEN	ANALYSIS TYPE	SOLVER TYPE	NODES	ELEMENTS
Type 1	Compression	LS-DYNA	285,521	843,996
Type 1	Three Point Bending	Static Structural	936,722	504,232
Type 2	Compression	LS-DYNA	277,530	833,096
Type 2	Three Point Bending	Static Structural	920,972	493,922

The numerical analysis of the compression tests was performed in ANSYS Workbench 2021 R1 using the LS-DYNA explicit solver, as shown in Figure 5. The LS-DYNA explicit solver has also been employed in recent studies to simulate the quasi-static compression and progressive deformation behavior of lattice structures [³⁴, ³⁵]. In the present study, the explicit formulation was preferred to accommodate the large deformations and complex contact interactions occurring during compression.

Figure 5
Finite element model of the compression analysis performed in the LS-DYNA environment.

The bottom surface of the specimen was fixed, and a total displacement of 24 mm was applied to the upper plate along the loading direction. A mesh size of 0.5 mm was used with tetrahedron elements to generate the mesh. During the analysis, the ratio of kinetic energy to internal energy was maintained below the commonly accepted threshold of 5% to ensure quasi-static conditions [³⁶]. Force–displacement curves were generated using the obtained reaction force data and compared with the experimental results.

A mesh convergence study was performed to verify that the key variables within the context of this study were not significantly affected by mesh density. Similar approaches have been reported in the literature [³⁷]. The mesh is considered converged when the variation between the results obtained with increasing mesh density is less than 5% [³⁸]. As shown in Figure 6, the results become stable after approximately 800,000 elements. In addition, the difference between the results obtained with finer meshes and the selected mesh was found to be limited. Therefore, a mesh corresponding to this level was adopted in the analyses in order to reduce computational cost.

Figure 6
Mesh convergence.

3. RESULTS AND DISCUSSION

As a result of the compression tests, local fracture and buckling behavior were observed in the struts aligned with the loading direction in the Z-oriented specimens, as illustrated in Figure 7. As the cells began to collapse, a distinct plateau region appeared in the load–displacement curve. Upon transition to the densification stage, where the cells fully collapsed, a significant increase in force was observed. The tests were terminated when the specimen reached approximately 50% deformation.

Figure 7
Deformation states of the Type 1 specimen: (a) 12 mm (25%) displacement and (b) 24 mm (50%) displacement.

As shown in Figure 8, it is seen that, in the Y-direction samples, deformation and buckling of the cells occurred under compression. An increase in displacement led to a reduction in the size of the cell voids.

Figure 8
Deformation states of the Type 2 specimen: (a) 12 mm (25%) displacement and (b) 24 mm (50%) displacement.

As the struts along the Y-axis were not aligned with the direction of loading, they offered no significant resistance to vertical loads. Consequently, structural deformation was observed at lower load magnitudes, leading to a lower plateau stress on the force-displacement plot.

According to the test results, as shown in Figure 9, specimens with vertical struts (Z-oriented) exhibited higher load-carrying capacity and more limited deformation at the same displacement level. The energy calculations were performed based on a displacement limit of approximately 20 mm, corresponding to the observed transition toward densification in the force–displacement curve. Beyond this point, the curve characteristics change due to the collapse of the cells; therefore, the energy evaluation was carried out up to 20 mm. The total absorbed energy is defined as the area under the load–displacement curve up to the onset of densification and is calculated using Equation (1) [³⁹].

Figure 9
Force–displacement curves obtained from the experimental and numerical compression tests of Type 1 and Type 2 specimens.

(1) $$E=\underset{0}{\overset{\text{u}}{{\displaystyle \int }}}P\left(x\right)dx$$

The total energy absorbed was computed using the force versus displacement results gathered from the compression tests. This was achieved by numerically integrating the area under the force-displacement curve using the trapezoidal rule. According to various literature, the energy dissipation capability of structures subjected to compression is evaluated based on Equation (2) [⁴⁰].

(2) $$\text{Ea}={\displaystyle \sum }_{i=2}^{n}0.5\left(P_i+P_{i-1}\right)\left({\delta }_i-{\delta }_{i-1}\right)$$

Specific Energy Absorption (SEA) is the amount of energy absorbed per unit weight during crushing distance and is an indication of the efficiency of energy utilization. Specific energy absorption can be determined from Equation (3) [⁴¹].

(3) $$SEA=\frac{E}{m}=\underset{0}{\overset{\text{u}}{{\displaystyle \int }}}\frac{P\left(x\right)dx}{m}$$

In the numerical analyses, SEA values were calculated using the ideal lattice structure mass obtained from the CAD model. In the experimental study, the measured actual mass of the fabricated specimens was used in the calculation of specific energy absorption, as presented in Table 3.

Table 3
Absorbed energy and specific energy absorption values of Type 1 and Type 2 lattice structures.

SPECIMEN	TOTAL ABSORBED ENERGY (J)	SEA (J/G)	EQUIVALENT ELASTIC MODULUS (MPa)	C1
Type 1 (Experimental)	22.009 ± 0.052	1.841	26.735	0.483
Type 1 (Numerical)	26.147	2.471	48.742	0.881
Type 2 (Experimental)	16.802 ± 0.690	1.405	23.179	0.419
Type 2 (Numerical)	19.600	1.852	41.083	0.742

The average energy absorption capacity of the vertically oriented specimens was found to be 22.009 J, whereas it was 16.802 J for the horizontally oriented specimens. Z-oriented (Type 1) specimens exhibited approximately 30.9% higher energy absorption compared to Y-oriented (Type 2) specimens. A similar trend was observed in the numerical analysis results, where Z-oriented structures demonstrated approximately 33.4% higher energy absorption than Y-oriented structures. Similar effects of lattice configuration on compressive energy absorption have been reported for SLA-fabricated polymer lattice structures, where variations in cell geometry and strut arrangement resulted in considerable differences in energy absorption performance [⁴²].

As shown in Table 4, a sensitivity analysis was performed to evaluate the effect of the selected cut-off displacement on the energy absorption results. The sensitivity analysis showed that the relative difference in absorbed energy between Type 1 and Type 2 remained similar at cut-off displacements of 18, 20, and 22 mm, ranging from 30.1% to 31.8% experimentally and from 32.0% to 33.9% numerically. Therefore, the 20 mm cut-off was retained for the energy absorption comparison.

Table 4
Sensitivity of absorbed energy to the selected cut-off displacement.

CUT-OFF DISPLACEMENT (mm)	TYPE 1 EXPERIMENTAL (J)	TYPE 1 NUMERICAL (J)	TYPE 2 EXPERIMENTAL (J)	TYPE 2 NUMERICAL (J)
18	18.654 ± 0.121	22.279	14.336 ± 0.454	16.883
20	22.009 ± 0.052	26.147	16.802 ± 0.690	19.600
22	26.112 ± 0.069	31.377	19.819 ± 0.780	23.434

The elastic modulus is defined as the resistance to deformation within the linear elastic region and is widely used to characterize the mechanical behavior of lattice structures. In calculating the equivalent elastic modulus, the lattice structure is assumed to behave as a continuum [²⁶]. The equivalent elastic modulus was determined from the linear region of the force–displacement curves. As shown in Table 3, Type 1 specimens exhibit higher equivalent elastic modulus values than Type 2 specimens in both experimental and numerical results. According to the experimental findings, the equivalent elastic modulus was 26.735 MPa for Type 1 specimens and 23.179 MPa for Type 2 specimens. In the numerical analyses, the equivalent elastic modulus values were calculated as 48.742 MPa and 41.083 MPa for Type 1 and Type 2 structures, respectively. The higher numerical equivalent elastic modulus resulted from the steeper initial slope of the numerical stress–strain curves compared with the experimental curves. A similar tendency, with steeper initial numerical slopes, can also be observed in a previous experimental–numerical study on MSLA lattice structures [¹³]. Manufacturing-related geometric imperfections in joints, corners, and struts may additionally contribute to such deviations [⁴³].

The elastic modulus of cellular structures is defined based on the Gibson–Ashby model as a function of relative density and the geometric constant C₁, and is calculated using Equation (4).

(4) $$E_l=C_1{E}_s{\left(\frac{{\mathrm{\rho }}_1}{{\mathrm{\rho }}_{\text{s}}}\right)}^{\text{n}}$$

For the geometric proportionality constant C1, Gibson and Ashby reported that this value can vary within the range of 0.1–4.0 [⁴⁴]. Here, Es represents the elastic modulus of the solid material. The ratio corresponds to the relative density, which is 0.3327 in this study. Gibson and Ashby demonstrated that n = 1 corresponds to stretching-dominated deformation mechanisms, while n = 2 is associated with bending-dominated mechanical behavior [⁴⁵]. In the present study, considering the Maxwell criterion, the structure was identified as bending-dominated, and therefore n was taken as 2. Based on the evaluated elastic modulus values, the C₁ coefficients were obtained as 0.881 for Type 1 and 0.742 for Type 2 in the numerical analyses, while the experimental values were 0.483 for Type 1 and 0.419 for Type 2.

As shown in Figures 10 and 11, the deformation behaviors of Type 1 and Type 2 structures show general similarities; however, the initiation regions and deformation mechanisms differ. In the Type 2 structure, crushing initiates from the top surface and progresses downward, whereas in the Type 1 structure, deformation begins from the bottom region. Additionally, in Type 1 structures, the struts aligned along the Z-direction expand outward under loading, resulting in noticeable lateral expansion. In contrast, Type 2 structures exhibit layer-by-layer crushing initiated from the top, accompanied by lateral deformation. These observations indicate that strut orientation directly influences the load transfer mechanism, local buckling regions, and overall deformation mode.

Figure 10
Comparison of experimental and numerical results of the Type 1 structure at (a) ε = 25% and (b) ε = 50% deformation.

Figure 11
Comparison of experimental and numerical results of the Type 2 structure at (a) ε = 25% and (b) ε = 50% deformation.

It was found that the point at which the two specimens began to approach the plateau zone was about 2.5 mm of displacement, while the transition toward densification was observed at approximately 20 mm displacement. As illustrated in Figure 12, the distribution of energy is mainly focused on the plateau zone. For instance, when the Type 1 specimen was subjected to experiments, the energy density in the elastic zone was determined to be 42.20 kJ/m³, but in the plateau zone, it was found to be 753.87 kJ/m³.

Figure 12
Comparison of elastic and plateau energy densities.

Based on the compression test findings, the influence of strut orientation on the deformation behavior and bending response of the lattice structures was further evaluated through three-point bending analyses.

After conducting the compression tests, the results of the three-point bending test analysis were considered for the study of flexural properties of the samples. As shown in Figure 13, a comparison of the experimental and numerical findings showed a strong correlation in the trends of the force-displacement plots.

Figure 13
Force–displacement curves obtained from experimental and numerical three-point bending tests of Type 1 and Type 2 specimens.

Stress concentrations have been found to occur in both the upper and lower portions of each configuration when comparing the images of both experimental and numerical. The Type 2 specimen has shown higher occurrences of buckling and deformation of struts, especially in the upper portion as can be seen from Figure 14 below.

Figure 14
Comparison of experimental and numerical deformations under three-point bending: (a) Type 2 and (b) Type 1 lattice structures.

The numerical results corresponding to approximately 10 mm displacement were compared with the equivalent experimental images, and it was determined that the deformation mechanisms are in good agreement. At a displacement level of 10 mm (Figure 15), the reaction force and absorbed energy for Z-oriented (Type 1) specimens were determined as approximately 60.65 N and 0.336 ± 0.004 J, respectively, whereas these values were 85.2 N and 0.471 ± 0.009 J for Y-oriented (Type 2) specimens. These results indicate that the Type 2 structure exhibits higher energy absorption capacity at the same displacement level. Previous studies on polymer lattice structures under three-point bending have similarly shown that changes in lattice configuration and geometry can result in substantial differences in load-carrying capacity and flexural response, highlighting the important role of structural configuration in bending behavior [⁴⁶, ⁴⁷]. A one-way analysis of variance (ANOVA) was performed to statistically evaluate the effect of lattice orientation on the energy absorption results. The differences between Type 1 and Type 2 specimens were statistically significant for both compression and three-point bending (p < 0.001).

Figure 15
Experimental images from three-point bending tests: (a) Type 1 and (b) Type 2 specimens.

4. CONCLUSIONS

This study examines the mechanical properties of both FBCCZ and FBCCY lattices with equal relative densities using experimental and computational methods. Experimental results reveal that the FBCCZ lattice, characterized by the presence of vertical strut members aligned with the applied loads, is capable of absorbing more energy than the FBCCY lattice.

This indicates the contribution of vertically aligned struts to the energy dissipation performance of the structure under axial loading.

The force–displacement curves indicate that both structures exhibit similar deformation behavior. In particular, the plateau region and the onset of densification occur at similar displacement levels. However, the higher slope observed in the pre-plateau region of the numerical curves leads to higher equivalent elastic modulus values in the numerical analyses.

In this study, the C1 coefficient was determined using the Gibson–Ashby model for lattice structures with different orientations. The C1 coefficient was calculated based on the experimentally obtained elastic modulus values. It is reported in the literature that this coefficient may vary depending on cell geometry, even at the same relative density [⁴⁴].

Layer-by-layer crushing behavior was observed in both structures; however, the initiation region of deformation differed depending on strut orientation. In FBCCY structures, crushing initiated from the top, whereas in FBCCZ structures, deformation began from the bottom.

It was also found that the force values obtained from finite element analyses were higher than those measured experimentally. This discrepancy is attributed to the assumption of ideal geometry and homogeneous material properties in the numerical model, as well as the exclusion of geometric imperfections and material defects arising during the manufacturing process.

Even after altering the specimen’s orientation, no significant changes in the energy distribution were observed. The study results showed that energy absorption occurs in the plateau zone regardless of orientation. Furthermore, the experimental findings were found to be in good agreement with the numerical analyses.

In the three-point bending tests, the FBCCY structure exhibited a higher load-carrying capacity compared to the FBCCZ structure. In the FBCCY structure, local buckling of the strut elements was observed in the upper regions near the loading point. It was observed that the horizontal elements provided more effective load transfer under bending. These findings indicate that the orientation of the strut elements relative to the loading direction has a significant effect on the bending behavior.

These results show that strut orientation, along with unit cell geometry, significantly affects load-bearing and deformation behavior. Accordingly, the optimal strut orientation depends on the loading condition. Z-oriented structures are more suitable for applications requiring high energy absorption under compressive loading, whereas Y-oriented structures provide superior performance under bending loads. Therefore, strut orientation should be considered as a key design parameter in the optimization of lattice structures for multi-loading applications.

5. DATA AVAILABILITY

The dataset supporting the results of this study is not publicly available.

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Corresponding Author:

Mustafa Arıkan, [e-mail], [e-mail], [e-mail]

Publication Dates

• Publication in this collection
  26 Oct 2026
• Date of issue
  2026

History

• Received
  19 May 2026
• Accepted
  01 Sept 2026