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29 pages, 9571 KiB  
Article
Design and Multi-Objective Optimization of Auxetic Sandwich Panels for Blastworthy Structures Using Machine Learning Method
by Andika, Sigit Puji Santosa, Djarot Widagdo and Arief Nur Pratomo
Appl. Sci. 2024, 14(23), 10831; https://doi.org/10.3390/app142310831 - 22 Nov 2024
Viewed by 402
Abstract
The design and multi-objective optimization of auxetic sandwich panels (ASPs) are performed to enhance the blastworthiness of armored fighting vehicles (AFVs). Various metastructures in the form of four auxetic geometries are proposed as the sandwich core: re-entrant honeycomb (REH), double-arrow honeycomb (DAH), star [...] Read more.
The design and multi-objective optimization of auxetic sandwich panels (ASPs) are performed to enhance the blastworthiness of armored fighting vehicles (AFVs). Various metastructures in the form of four auxetic geometries are proposed as the sandwich core: re-entrant honeycomb (REH), double-arrow honeycomb (DAH), star honeycomb (SH), and tetra-chiral honeycomb (CH). This paper employs a combination of finite element and machine learning methodologies to evaluate blastworthiness performance. Optimization is carried out using the nondominated sorting genetic algorithm II (NSGA-II) method. The optimization results show significant improvements in blastworthiness performance, with notable reductions in permanent displacement and enhancements in specific energy absorption (SEA). Global sensitivity analysis using SHapley Additive exPlanations (SHAP) reveals that cell thickness is the most critical factor affecting blastworthiness performance, followed by the number of cells and corner angle or radius for CH. The application of optimized ASP on AFVs shows promising results, with no failure occurring in the occupant floor. Furthermore, AFVs equipped with the optimized ASP DAH significantly reduce maximum displacement and acceleration by 39.00% and 43.56%, respectively, and enhance SEA by 48.30% compared to optimized aluminum foam sandwich panels. This study concludes that ASPs have potential applications in broader engineering fields. Full article
(This article belongs to the Special Issue Structural Dynamics and Protective Materials)
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Figure 1
<p>(<b>a</b>) Configuration of the auxetic sandwich panel (ASP), which consists of two face sheets and an auxetic core. (<b>b</b>) Overall dimensions of the ASP: <math display="inline"><semantics> <mrow> <mi>L</mi> <mo>×</mo> <mi>W</mi> <mo>×</mo> <mi>H</mi> </mrow> </semantics></math>. (<b>c</b>) Zoomed-in view of the face sheets and auxetic core of the ASP, highlighting the region marked by the red dashed line in panel (<b>b</b>).</p>
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<p>Geometric parameters of the unit cell for (<b>a</b>) re-entrant honeycomb (REH), (<b>b</b>) double-arrow honeycomb (DAH), (<b>c</b>) star honeycomb (SH), and (<b>d</b>) tetra-chiral honeycomb (CH).</p>
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<p>Finite element (FE) model of the ASP subjected to air-blast loading at the center. The ASP is modeled with quarter symmetry due to its symmetric response. The red, green, and blue lines represent the fixed boundary condition and the symmetric boundary conditions in the <span class="html-italic">y</span>–<span class="html-italic">z</span> and <span class="html-italic">x</span>–<span class="html-italic">z</span> planes, respectively.</p>
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<p>The curve fitting of modified Johnson–Cook parameters of 304 stainless steel with the experiment [<a href="#B14-applsci-14-10831" class="html-bibr">14</a>].</p>
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<p>Quarter-symmetric model of a single plate subjected to air-blast loading using three types of air-blast models: (<b>a</b>) Conventional Weapons Effects Program (CONWEP), (<b>b</b>) smoothed-particle hydrodynamics (SPH), and (<b>c</b>) multi-material Arbitrary Lagrangian–Eulerian (MMALE). In the SPH model, the black line represents the SPH calculation domain. In the MMALE model, the C-4 charge and air are modeled using Eulerian formulations, while the single plate is modeled using Lagrangian formulations.</p>
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<p>Comparison of permanent displacement (<math display="inline"><semantics> <msub> <mi>δ</mi> <mi>p</mi> </msub> </semantics></math>) from various air-blast models with experimental results [<a href="#B6-applsci-14-10831" class="html-bibr">6</a>].</p>
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<p>(<b>a</b>) Experimental setup for the blast test of trapezoidal corrugated-core sandwich panels. (<b>b</b>) Configuration of the trapezoidal corrugated-core sandwich panel [<a href="#B14-applsci-14-10831" class="html-bibr">14</a>].</p>
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<p>Comparison of permanent displacement (<math display="inline"><semantics> <msub> <mi>δ</mi> <mi>p</mi> </msub> </semantics></math>) between numerical and experimental results [<a href="#B14-applsci-14-10831" class="html-bibr">14</a>] for trapezoidal corrugated-core sandwich panels.</p>
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<p>Comparison of deformation (<b>left</b>) and cross-sectional displacement profiles of face sheets (<b>right</b>) between numerical simulations and experimental results [<a href="#B14-applsci-14-10831" class="html-bibr">14</a>] for specimens (<b>a</b>) TZ-12, (<b>b</b>) TZ-2, and (<b>c</b>) TZ-13. The red and green dashed lines represent plastic buckling and local indentation after blast impact, respectively.</p>
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<p>Design space of (<b>a</b>) REH, (<b>b</b>) DAH, (<b>c</b>) SH, and (<b>d</b>) CH for multi-objective optimization problem (MOOP). The color represents the normalized relative density (<math display="inline"><semantics> <msub> <mi>ρ</mi> <mi>r</mi> </msub> </semantics></math>).</p>
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<p>Scatter plots of predicted values from metamodel versus simulation results of <math display="inline"><semantics> <msub> <mi>δ</mi> <mi>p</mi> </msub> </semantics></math> (<b>left</b>) and SEA (<b>right</b>) for (<b>a</b>) REH, (<b>b</b>) DAH, (<b>c</b>) SH, and (<b>d</b>) CH models. The black diagonal line represents the ideal case where predicted values match simulation values, indicating perfect accuracy.</p>
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<p>Averaged SHapley Additive exPlanations (SHAP) values of (<b>a</b>) <math display="inline"><semantics> <msub> <mi>δ</mi> <mi>p</mi> </msub> </semantics></math> and (<b>b</b>) SEA for each auxetic model.</p>
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<p>SHAP summary plot of <math display="inline"><semantics> <msub> <mi>δ</mi> <mi>p</mi> </msub> </semantics></math> (<b>left</b>) and SEA (<b>right</b>) for (<b>a</b>) REH, (<b>b</b>) DAH, (<b>c</b>) SH, and (<b>d</b>) CH models.</p>
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<p>Scatter plot showing the influence of input variables for (<b>a</b>) REH, (<b>b</b>) DAH, (<b>c</b>) SH, and (<b>d</b>) CH models.</p>
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<p>Pareto front of the MOOP for (<b>a</b>) REH, (<b>b</b>) DAH, (<b>c</b>) SH, and (<b>d</b>) CH models. The Pareto front is limited to <math display="inline"><semantics> <mrow> <msub> <mi>δ</mi> <mi>p</mi> </msub> <mo>&lt;</mo> <mn>50</mn> </mrow> </semantics></math> mm.</p>
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<p>Comparison of optimized ASP models: (<b>a</b>) REH, (<b>b</b>) DAH, (<b>c</b>) SH, (<b>d</b>) CH; and (<b>e</b>) optimized aluminum foam sandwich panel (AFSP) model [<a href="#B52-applsci-14-10831" class="html-bibr">52</a>].</p>
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<p>Comparison of the Pareto front for all auxetic models.</p>
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<p>Deformed shapes of four model types for each ASP, (<b>a</b>) REH, (<b>b</b>) DAH, (<b>c</b>) SH, and (<b>d</b>) CH, at 5 ms. The configurations of each model type are listed in <a href="#applsci-14-10831-t008" class="html-table">Table 8</a>. From left (lowest relative density) to right (highest relative density): ideal design of maximum SEA, balanced design, baseline design, and ideal design of minimum <math display="inline"><semantics> <msub> <mi>δ</mi> <mi>p</mi> </msub> </semantics></math>.</p>
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<p>Displacement vectors of ASP: (<b>a</b>) REH, (<b>b</b>) DAH, (<b>c</b>) SH, and (<b>d</b>) CH. The red lines indicate the areas that commonly exhibit negative Poisson’s ratio (NPR) behavior except for the SH model.</p>
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<p>The FE model of the armored fighting vehicle (AFV) subsystem under 8 kg TNT: (<b>a</b>) side view, (<b>b</b>) front view, (<b>c</b>) detailed view of 8 kg cylindrical TNT placed on a rigid steel pot.</p>
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<p>Failure modes of the occupant and struck sides (<b>top</b>) and deformation process from 0 to 10 ms (<b>bottom</b>) of the AFV subsystem for different types of structures: (<b>a</b>) 10 mm occupant side plate (OSP) without a protective structure, (<b>b</b>) 20 mm OSP without a protective structure, (<b>c</b>) with optimized AFSP, (<b>d</b>) with optimized ASP DAH.</p>
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<p>Dynamic responses of (<b>a</b>) displacement and (<b>b</b>) acceleration on the occupant side of the AFV for different structure types.</p>
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<p>Comparison of (<b>a</b>) energy absorption (EA) and (<b>b</b>) SEA for each part of the AFV subsystem for different structure types.</p>
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18 pages, 3755 KiB  
Article
Immunomodulatory Effects of the Tobacco Defensin NaD1
by Ekaterina I. Finkina, Ivan V. Bogdanov, Olga V. Shevchenko, Serafima I. Fateeva, Anastasia A. Ignatova, Sergey V. Balandin and Tatiana V. Ovchinnikova
Antibiotics 2024, 13(11), 1101; https://doi.org/10.3390/antibiotics13111101 - 19 Nov 2024
Viewed by 404
Abstract
Background/Objectives: Defensins are important components of the innate plant immune system, exhibiting antimicrobial activity against phytopathogens, as well as against fungi pathogenic to humans. Along with antifungal activity, plant defensins are also capable of influencing various immune processes, but not much is known [...] Read more.
Background/Objectives: Defensins are important components of the innate plant immune system, exhibiting antimicrobial activity against phytopathogens, as well as against fungi pathogenic to humans. Along with antifungal activity, plant defensins are also capable of influencing various immune processes, but not much is known about these effects. In this study, we investigated the immunomodulatory effects of the tobacco defensin NaD1, which possesses a pronounced antifungal activity. Methods and Results: We showed that NaD1 could penetrate the Caco-2 polarized monolayer. Using a multiplex assay with a panel of 48 cytokines, chemokines and growth factors, we demonstrated that NaD1 at a concentration of 2 μM had immunomodulatory effects on human dendritic cells and blood monocytes, mainly inhibiting the production of various immune factors. Using the sandwich ELISA method, we demonstrated that NaD1 at the same concentration had a pronounced immunomodulatory effect on unstimulated THP-1-derived macrophages and those stimulated by bacterial LPS or fungal zymosan. NaD1 had a dual effect and induced the production of both pro-inflammatory cytokine IL-1β as well as anti-inflammatory IL-10 on resting and pro-inflammatory THP-1-derived macrophages. We also found that the immunomodulatory effects of the tobacco defensin NaD1 and the pea defensin Psd1 differed from each other, indicating nonuniformity in the modes of action of plant defensins. Conclusions: Thus, our data demonstrated that the tobacco defensin NaD1 exhibits different immunomodulatory effects on various immune cells. We hypothesized that influence on human immune system along with antifungal activity, could determine the effectiveness of this peptide under infection in vivo. Full article
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<p>Cytotoxic effects of the tobacco defensin NaD1 towards PBMCs (<b>A</b>) and Caco-2 cells in monolayer (<b>C</b>). The membrane-active peptide melittin from the venom of honeybees (<b>B</b>,<b>D</b>) was used for comparison. Error bars represent a standard deviation (±SD) between two biological and two technical replications. Significance levels are * <span class="html-italic">p</span> ≤ 0.05, *** <span class="html-italic">p</span> &lt; 0.001 and **** <span class="html-italic">p</span> &lt; 0.0001. The significance was calculated by comparing untreated cells (control) with treated by NaD1 or melittin cells. Viability cells in control and experimental samples was compared with un-paired two-sample <span class="html-italic">t</span>-test.</p>
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<p>Assessment of bidirectional transport of the tobacco defensin NaD1 through the polarized Caco-2 monolayer. A→B, absorptive transport; B→A, secretory transport; Papp—apparent permeability coefficient. Six and four independent biological replications were used for absorptive and secretory directions, respectively. The normality of Papp coefficient distribution was assessed using Shapiro–Wilk test. Papp coefficients were compared by unpaired two-sample <span class="html-italic">t</span>-test.</p>
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<p>Production of cytokines, chemokines and growth factors upon stimulation of DCs and monocytes by NaD1 at the concentration of 2 μM. Error bars represent a standard deviation (±SD) between two biological replications. The levels in control and experimental wells were compared by unpaired two-sample <span class="html-italic">t</span>-test.</p>
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<p>Influence of the tobacco defensin NaD1 and other AMPs at the concentration of 2 µM on production of pro- (<b>A</b>–<b>D</b>) and anti-inflammatory (<b>B</b>) cytokines either unstimulated or stimulated by LPS or by zymosan THP-1-derived macrophages. Error bars represent a standard deviation (±SD) between two biological and two technical replications. Significance levels are * <span class="html-italic">p</span> ≤ 0.05, ** <span class="html-italic">p</span> &lt; 0.01, *** <span class="html-italic">p</span> &lt; 0.001, **** <span class="html-italic">p</span> &lt; 0.0001. The significance of difference in cytokine production was calculated by comparing: unstimulated cells (control) with stimulated by AMPs cells (grey bars); stimulated by LPS (blue bars) or zymosan (green bars) cells alone or in the presence of AMPs. Release of the cytokines in control and experimental samples was compared with unpaired two-sample <span class="html-italic">t</span>-test.</p>
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21 pages, 18918 KiB  
Article
Structural and Sustainability Enhancement of Composite Sandwich Slab Panels Using Novel Fibre-Reinforced Geopolymer Concrete
by M. Sridhar and M. Vinod Kumar
J. Compos. Sci. 2024, 8(11), 479; https://doi.org/10.3390/jcs8110479 - 18 Nov 2024
Viewed by 379
Abstract
One of the important findings of the recent decades in the construction industry is composite sandwich panels (CSPs), which have benefits of being lightweight, providing thermal insulation, and aiding the economy; they are transforming continuously through many add-ons as needed by the industry. [...] Read more.
One of the important findings of the recent decades in the construction industry is composite sandwich panels (CSPs), which have benefits of being lightweight, providing thermal insulation, and aiding the economy; they are transforming continuously through many add-ons as needed by the industry. With the demand for sustainability in the field, CSPs need structural and sustainable enhancement. In the present study, an approach for the same has been attempted with geopolymer concrete (GPC) and novel nylon fibre to improve the sustainability and structural benefits, respectively. With various material combinations including GPC reinforced with fibres, six CSPs were cast and studied. The inherent limitations of GPC have been addressed by the nylon fibre reinforcement instead of using steel fibres, which have a similar strength, considering the aim of maintaining the density of the wythe material. A comparison of the flexural behaviour of the CSPs through the parameters of load–deflection, ductility, and toughness was made using the four-point loading test. The results of the test specify that the fibres enhance the performance of the CSPs under flexural loading. Full article
(This article belongs to the Section Composites Applications)
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<p>An illustration of the effect of the FENF over straight fibres under flexure: (<b>a</b>) a member subjected to flexural loading, (<b>b</b>) a member without fibre reinforcement with concentrated cracks and wide cracks, (<b>c</b>) a member with nylon fibre reinforcement with distributed cracks and reduced crack width, and (<b>d</b>) a member with FENF reinforcement (many distributed cracks and fine crack width).</p>
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<p>Work breakdown and methodology.</p>
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<p>Illustration of FENF preparation: (<b>a</b>) raw nylon fibres before modification, (<b>b</b>) straight nylon fibres cut to specified lengths, (<b>c</b>) collection of FENFs after end-flattening, and (<b>d</b>) enlarged view of FENFs, showing straight portion and flattened ends.</p>
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<p>Dimensions of EPS panel and reinforcement wire meshes.</p>
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<p>Casting process of CSPs: (<b>a</b>) EPS panels placed in casting yard; (<b>b</b>) top wythe concreting; (<b>c</b>) flipping panels for bottom wythe concreting; (<b>d</b>) final cast panels.</p>
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<p>Schematic loading and instrumentation setup.</p>
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<p>Loading arrangement with data acquisition system.</p>
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<p>Load vs. deflection curves for all panel types.</p>
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<p>Yield and ultimate load capacity.</p>
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<p>Deflection at yield and ultimate loads.</p>
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<p>Ductility.</p>
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<p>Initial stiffness.</p>
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<p>Crack width development vs. applied load.</p>
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<p>Crack width measured at yield loading during testing: (<b>a</b>) CC, (<b>b</b>) CG, (<b>c</b>) GG, (<b>d</b>) CGSN, (<b>e</b>) CGFN, and (<b>f</b>) CGHS.</p>
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<p>Bar chart comparing toughness and energy dissipation.</p>
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20 pages, 11339 KiB  
Article
Modeling the Dynamic Properties of Multi-Layer Glass Fabric Sandwich Panels
by Arkadiusz Charuk, Izabela Irska and Paweł Dunaj
Polymers 2024, 16(21), 3074; https://doi.org/10.3390/polym16213074 - 31 Oct 2024
Viewed by 513
Abstract
Sandwich panels are key components of many lightweight structures. They are often subjected to time-varying loads, which can cause various types of vibrations that adversely affect the functionality of the structure. That is why it is of such importance to predict the dynamic [...] Read more.
Sandwich panels are key components of many lightweight structures. They are often subjected to time-varying loads, which can cause various types of vibrations that adversely affect the functionality of the structure. That is why it is of such importance to predict the dynamic properties of both the panels and the structures made of them at the design stage. This paper presents finite element modeling of the dynamic properties (i.e., natural frequencies, mode shapes, and frequency response functions) of sandwich panels made of glass fabric impregnated with phenolic resin. The model reproducing the details of the panel structure was built using two-dimensional, quadrilateral, isoparametric plane elements. Afterwards, the model was subjected to an updating procedure based on experimentally determined frequency response functions. As a result, the average relative error for natural frequencies achieved numerically was 5.0%. Finally, a cabinet model consisting of the analyzed panels was built and experimentally verified. The relative error between the numerically and experimentally obtained natural frequencies was on average 5.9%. Full article
(This article belongs to the Special Issue Damping Mechanisms in Polymers and Polymer Composites)
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<p>Structure of a multi-layer glass fabric sandwich panel—general view.</p>
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<p>The cross-sectional optical images of the analyzed panel: parallel to the tunnel alignment (<b>a</b>,<b>b</b>), perpendicular to the tunnel alignment (<b>c</b>), and hollow tunnel geometry (zoomed view) (<b>d</b>).</p>
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<p>Representative stress-strain curves of top, internal, and bottom plates of the analyzed panel tested perpendicular (<b>a</b>) and parallel (<b>b</b>) to the z-reinforcement. Summary of tensile characterization (<b>c</b>).</p>
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<p>Representative stress-strain curves of top, internal, and bottom plates of the analyzed panel tested perpendicular (<b>a</b>) and parallel (<b>b</b>) to the z-reinforcement; summary of flexural characterization (<b>c</b>).</p>
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<p>Finite element model of the analyzed panel.</p>
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<p>Modal analysis test stand with measurement points arrangement: schematic representation (<b>a</b>) and actual stand (<b>b</b>).</p>
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<p>Comparison of selected mode shapes determined numerically and experimentally.</p>
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<p>Comparison of selected frequency response functions determined numerically and experimentally.</p>
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<p>Sensitivity analysis results for analyzed panel.</p>
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<p>Comparison of selected frequency response functions determined numerically (before and after model updating) and experimentally.</p>
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<p>Comparison of selected frequency response functions determined numerically and experimentally for 390 × 796 × 18 mm panel.</p>
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<p>A finite element model of cabinet.</p>
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<p>Modal analysis test stand—cabinet testing.</p>
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<p>Comparison of cabinet mode shapes determined numerically and experimentally.</p>
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<p>Comparison of frequency response functions for cabinet determined numerically and experimentally.</p>
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16 pages, 4859 KiB  
Article
Study on the Dynamic Crushing Behaviors of Hourglass Honeycomb Sandwich Panels
by Xinhai Chen, Kai Wang, Lu Cao, Pengyu Guo, Jiangyi Qin and Hexiang Wu
Aerospace 2024, 11(11), 881; https://doi.org/10.3390/aerospace11110881 - 25 Oct 2024
Viewed by 520
Abstract
In response to the problem of enclosed internal spaces in existing honeycomb sandwich panels, the concept of an hourglass honeycomb sandwich panel model is proposed for the first time, which provides a breakthrough approach for achieving the multifunctional integration of honeycomb sandwich panels. [...] Read more.
In response to the problem of enclosed internal spaces in existing honeycomb sandwich panels, the concept of an hourglass honeycomb sandwich panel model is proposed for the first time, which provides a breakthrough approach for achieving the multifunctional integration of honeycomb sandwich panels. Numerical simulation methods are employed to investigate the dynamic performance of the hourglass honeycomb sandwich panels. The focus is on discussing the influences of the geometric parameters on the deformation mode, dynamic response, load uniformity, and energy absorption capacity of the hourglass honeycomb sandwich panel under different impact velocity conditions. The research results indicate that under low-velocity-impact conditions, the influence of the geometric parameters is predominant. In contrast, under high-velocity-impact conditions, the influence of the impact velocity conditions is predominant. Hourglass honeycomb sandwich panels with low density, a large inclination angle of the honeycomb wall, and small contact distances between the hourglass honeycomb cell and the panel have excellent load uniformity, and the distances between the contact points of the hourglass honeycomb cell and the panel have a great influence on the energy absorption capacity of the sandwich panels. This study provides a theoretical basis for the application of honeycombs in aerospace and other engineering areas. Full article
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<p>Computational model of hourglass honeycomb sandwich panel.</p>
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<p>Geometric model: (<b>a</b>) hourglass honeycombs; (<b>b</b>) unit cell configuration.</p>
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<p>Comparative validation of experimental results and simulated results for sandwich panels: (<b>a</b>) static experimental results [<a href="#B43-aerospace-11-00881" class="html-bibr">43</a>]; (<b>b</b>) dynamic experimental results [<a href="#B44-aerospace-11-00881" class="html-bibr">44</a>].</p>
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<p>Deformation patterns at <span class="html-italic">v</span> = 2 m/s: (<b>a</b>) circular honeycomb sandwich panel (H0); (<b>b</b>) hourglass honeycomb sandwich panel (H1).</p>
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<p>Deformation patterns at <span class="html-italic">v</span> = 30 m/s: (<b>a</b>) circular honeycomb sandwich panel (H0); (<b>b</b>) hourglass honeycomb sandwich panel (H1).</p>
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<p>Deformation patterns at <span class="html-italic">v</span> = 120 m/s: (<b>a</b>) circular honeycomb sandwich panel (H0); (<b>b</b>) hourglass honeycomb sandwich panel (H1).</p>
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<p>Nominal stress–strain curves of honeycomb sandwich panels at <span class="html-italic">v</span> = 2 m/s: (<b>a</b>) different honeycomb wall thicknesses; (<b>b</b>) different honeycomb wall inclination angles; (<b>c</b>) different maximum cell radii; (<b>d</b>) different numbers of honeycomb core layers.</p>
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<p>Nominal stress–strain curves of honeycomb sandwich panels at <span class="html-italic">v</span> = 30 m/s: (<b>a</b>) different honeycomb wall thicknesses; (<b>b</b>) different honeycomb wall inclination angles; (<b>c</b>) different maximum cell radii; (<b>d</b>) different numbers of honeycomb core layers.</p>
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<p>Nominal stress–strain curves of honeycomb sandwich panels at <span class="html-italic">v</span> = 120 m/s: (<b>a</b>) different honeycomb wall thicknesses; (<b>b</b>) different honeycomb wall inclination angles; (<b>c</b>) different maximum cell radii; (<b>d</b>) different numbers of honeycomb core layers.</p>
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<p>Nominal stress–strain curves of honeycombs.</p>
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<p>The load uniformity of honeycomb sandwich panels under different impact velocities: (<b>a</b>) <span class="html-italic">v</span> = 2 m/s; (<b>b</b>) <span class="html-italic">v</span> = 30 m/s; (<b>c</b>) <span class="html-italic">v</span> = 120 m/s.</p>
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<p>The unit mass energy absorption curves of honeycomb sandwich panels under different impact conditions: (<b>a</b>) <span class="html-italic">v</span> = 2 m/s; (<b>b</b>) <span class="html-italic">v</span> = 30 m/s; (<b>c</b>) <span class="html-italic">v</span> = 120 m/s.</p>
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15 pages, 4925 KiB  
Article
The Evaluation of Sandwich Composite Materials with Vegetable Fibers in a Castor Oil Polyurethane Matrix with Their Faces and Honeycomb Core Made in a 3D Printer
by Gilberto Garcia del Pino, Abderrezak Bezazi, Antonio Claudio Kieling, José Costa de Macedo Neto, Sofia Dehaini Garcia, José Luis Valin Rivera, Meylí Valin Fernández, Aristides Rivera Torres and Francisco Rolando Valenzuela Diaz
Polymers 2024, 16(21), 2980; https://doi.org/10.3390/polym16212980 - 24 Oct 2024
Viewed by 581
Abstract
Sandwich panels are widely used in the naval and aerospace industries to withstand the normal tensile, compressive, and shear stresses associated with bending. The faces of sandwich composites are usually made of metals such as aluminum and, in some studies with composites, using [...] Read more.
Sandwich panels are widely used in the naval and aerospace industries to withstand the normal tensile, compressive, and shear stresses associated with bending. The faces of sandwich composites are usually made of metals such as aluminum and, in some studies with composites, using a polymeric matrix, but there are no studies in the literature using a castor oil polyurethane matrix. The core of the panel must keep the faces apart and be rigid perpendicular to them. To begin the work, a study was carried out on the influence of alkaline treatment on sisal fibers to increase the fibers’ adhesion to castor oil polyurethane. There are no relevant studies worldwide on the use of this resin and the adhesion of vegetable fibers to this polyurethane. In this work, a study was carried out through a three-point bending test of sandwich panels using faces of composite material with sisal fibers subjected to an alkaline treatment of 10% by weight of sodium hydroxide and an immersion time of 4 h in the dissolution, which was the best chemical treatment obtained initially in a castor oil polyurethane matrix. The honeycomb cores were made by 3D printer and in this study two different printing filament materials, PETG and PLA, and two different core heights were compared. As a result of a traction test, it was observed that sisal fibers with chemical treatment in a castor oil polyurethane matrix can be used in composites, although the stress levels obtained are 50% lower than the stresses obtained in other matrixes such as epoxy resin. The combination of sisal faces in a castor oil polyurethane matrix and honeycomb cores made in a 3D printer showed good properties, which allows the use of renewable, sustainable and less aggressive materials for the environment. In all tests, PETG was 21% to 32% stronger than PLA. Although there was no rupture in the test specimens, the PETG cores deformed 0.5% to 3.6% less than PLA. The composites with PLA were lighter, because the core density was 13.8% lower than the PETG cores. Increasing the height of the honeycomb increased its strength. Full article
(This article belongs to the Section Biobased and Biodegradable Polymers)
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<p>Sisal fiber processing in the State of Bahia, Brazil: (<b>a</b>) sisal planting, (<b>b</b>) collection, and (<b>c</b>) the drying of the fiber extracted from the leaves [<a href="#B13-polymers-16-02980" class="html-bibr">13</a>].</p>
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<p>Castor oil plant with fruits containing seeds [<a href="#B16-polymers-16-02980" class="html-bibr">16</a>].</p>
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<p>Manufacturing test specimens: (<b>a</b>) the chemical treatment of the fibers, (<b>b</b>) the placement of the fibers and resin inside the mold, (<b>c</b>) closing the mold, (<b>d</b>) the cutting of the test specimens with a laser.</p>
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<p>Tensile tests: (<b>a</b>) the start of the application of the load on the traction machine, (<b>b</b>) the end of the test at the moment of rupture.</p>
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<p>Manufacturing of the composite faces or sheet: (<b>a</b>) the alkaline treatment, (<b>b</b>) the mold for the manufacture of the sisal/epoxy composite sheets, (<b>c</b>) the placement of the fibers and resin inside the mold, (<b>d</b>) face length, (<b>e</b>) face width.</p>
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<p>Printer configuration: (<b>a</b>) hexagon dimensions, (<b>b</b>) honeycomb model, (<b>c</b>) CREALITY ENDER 3 V2 printer model.</p>
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<p>Printing parameters: (<b>a</b>) E–STEP calibration, (<b>b</b>) printing and (<b>c</b>) calibration cube measurement.</p>
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<p>The manufacturing and gluing of the honeycombs. (<b>a</b>) The printing of the honeycombs and (<b>b</b>) the gluing of the first side.</p>
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<p>Top face bonding.</p>
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<p>The graphs generated by the tensile testing machine for the test specimens with the highest tensile stress corresponding to an alkaline treatment with a 10% concentration and 4 h of immersion of the fiber in the dissolution.</p>
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<p>The influence of alkaline treatment on tensile stress: (<b>a</b>) the behavior of the average tensile stress as a function of immersion time, (<b>b</b>) the influence of NaOH concentration on tensile stress for an immersion time of 4 h.</p>
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<p>Three-point bending test of specimen 1.</p>
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<p>Three-point bending test of specimen 2.</p>
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<p>Bending load–displacement curves from the three-point bending tests for the 10 mm core height test specimens.</p>
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<p>Bending load–displacement curves from the three-point bending tests for the 15 mm core height test specimens.</p>
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<p>Specimens (1–4) After Bending Tests.</p>
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<p>Influence of material and honeycomb height on maximum bending stress.</p>
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21 pages, 3616 KiB  
Review
A Review of the Biomimetic Structural Design of Sandwich Composite Materials
by Shanlong Che, Guangliang Qu, Guochen Wang, Yunyan Hao, Jiao Sun and Jin Ding
Polymers 2024, 16(20), 2925; https://doi.org/10.3390/polym16202925 - 18 Oct 2024
Viewed by 1185
Abstract
Sandwich composites are widely used in engineering due to their excellent mechanical properties. Accordingly, the problem of interface bonding between their panels and core layers has always been a hot research topic. The emergence of biomimetic technology has enabled the integration of the [...] Read more.
Sandwich composites are widely used in engineering due to their excellent mechanical properties. Accordingly, the problem of interface bonding between their panels and core layers has always been a hot research topic. The emergence of biomimetic technology has enabled the integration of the structure and function of biological materials from living organisms or nature into the design of sandwich composites, greatly improving the interface bonding and overall performance of heterogeneous materials. In this paper, we review the most commonly used biomimetic structures and the fusion design of multi-biomimetic structures in the engineering field. They are analyzed with respect to their mechanical properties, and several biomimetic structures derived from abstraction in plants and animals are highlighted. Their structural advantages are further discussed specifically. Regarding the optimization of different interface combinations of multilayer composites, this paper explores the optimization of simulations and the contributions of molecular dynamics, machine learning, and other techniques used for optimization. Additionally, the latest molding methods for sandwich composites based on biomimetic structural design are introduced, and the materials applicable to different processes, as well as their advantages and disadvantages, are briefly analyzed. Our research results can help improve the mechanical properties of sandwich composites and promote the application of biomimetic structures in engineering. Full article
(This article belongs to the Section Polymer Composites and Nanocomposites)
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<p>Biomimetic structural design of shell nacre (<b>a</b>) imitation shell pearl layer structure preparation (<b>b</b>) soft material and specimen face in the direction of the angle (<b>c</b>) soft material and specimen face outside the direction of the angle [<a href="#B24-polymers-16-02925" class="html-bibr">24</a>].</p>
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<p>Biomimetic design of the surface structure and model of the mantis shrimp: (<b>a</b>) representative appearance of Odontodactylus scyllarus with two integrated dactyls; (<b>b</b>) individual drawing of the dactyl club; (<b>c</b>) CT scanning of a dactyl club section; (<b>d</b>) bio-inspired bi-directionally sinusoidal corrugated panel; (<b>e</b>) bio-inspired bi directionally sinusoidal corrugated sandwich structure [<a href="#B28-polymers-16-02925" class="html-bibr">28</a>].</p>
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<p>Detailed microstructure and geometric model of pomelo peel [<a href="#B38-polymers-16-02925" class="html-bibr">38</a>].</p>
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<p>Finite element models of traditional honeycomb and pomelo peel honeycomb [<a href="#B38-polymers-16-02925" class="html-bibr">38</a>].</p>
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<p>Biomimetic structural model of wood cell walls with different spiral angles [<a href="#B42-polymers-16-02925" class="html-bibr">42</a>].</p>
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<p>Biological prototypes and design strategies for the fusion of biomimetic designs. (<b>a</b>,<b>a1</b>) Schematic diagram of the S-shaped structure of cuttlefish bone. (<b>a2</b>) Bionic design of the S-shaped structure of cuttlefish bone. (<b>b</b>) Spider web, the second biological prototype used for integrated bionic design. (<b>b1</b>) Schematic diagram of the structure of spider web. (<b>b2</b>) Bionic structure of spider web. (<b>c</b>) Pomelo, the third biological prototype used for integrated bionic design. (<b>c1</b>) Schematic diagram of the porous structure of pomelo peel. (<b>c2</b>) Bionic design of the porous structure of pomelo peel. (<b>d1</b>) The integration of a spiderweb-like structure (<b>b2</b>) and a cuttlefish bone-like s-shaped edge (<b>a2</b>). (<b>d2</b>) Foam that imitates the porous structure of pomelo peel. (<b>e</b>) The S-shaped spider web integrated with porous foam to form a flexible composite of porous materials with an S-shaped spider web structure that was designed in this study [<a href="#B49-polymers-16-02925" class="html-bibr">49</a>].</p>
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<p>Finite element analysis results of impact mechanical properties of the integrated bio-inspired structure (<b>a</b>) FE simulation results of an unfilled contrast structure. (<b>b</b>) FE simulation results of an unfilled spider web. (<b>c</b>) FE simulation results of an unfilled S-spider web. (<b>d</b>) FE simulation results of a contrast structure-gel. (<b>e</b>) FE simulation results of a spider web-gel. (<b>f</b>) FE simulation results of an S-spider web-gel. (<b>g</b>) FE simulation results of a contrast structure-foam. (<b>h</b>) FE simulation results ofa spider web-foam. (<b>i</b>) FE simulation results of an S-spider web-foam [<a href="#B49-polymers-16-02925" class="html-bibr">49</a>].</p>
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<p>Design of angle-reinforced spider web honeycomb biomimetic structure. (<b>a</b>) Natural spiderweb structure and its simplified representation; (<b>b</b>) corner- enhanced biomimetic spiderweb-structured honeycomb models; and (<b>c</b>) corner-enhanced biomimetic spiderweb hierarchies [<a href="#B53-polymers-16-02925" class="html-bibr">53</a>].</p>
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<p>A review of the typical design of different biomimetic structures.</p>
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<p>Design of interlocked structure imitating the exoskeleton of the armored beetle (<b>a</b>) biological suture interface; (<b>b</b>) design of the specimen model; (<b>c</b>) parameters of inter locked tooth [<a href="#B75-polymers-16-02925" class="html-bibr">75</a>].</p>
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<p>(<b>a</b>–<b>c</b>) Low-power SEM images of MCMB@WC crack growth path under different volume fractions: (<b>a</b>) 17%; (<b>b</b>) 53%; (<b>c</b>) non-biomimetic structure; (<b>d</b>–<b>f</b>) high-power SEM images of MCMB@WC crack growth path under different volume fractions: (<b>d</b>) 17%; (<b>e</b>) 53%; (<b>f</b>) non-biomimetic structure; (<b>g</b>–<b>i</b>) schematic diagram of MCMB@WC crack growth under different volume fractions: (<b>g</b>) 17%; (<b>h</b>) 53% and (<b>i</b>) non-biomimetic structures [<a href="#B13-polymers-16-02925" class="html-bibr">13</a>].</p>
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<p>Finite element analysis of the bending performance of interlocked structures [<a href="#B86-polymers-16-02925" class="html-bibr">86</a>].</p>
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20 pages, 21682 KiB  
Article
Particularities on the Low-Velocity Impact Behavior of 3D-Printed Sandwich Panels with Re-Entrant and Honeycomb Core Topologies
by Andrei Ioan Indreș, Dan Mihai Constantinescu, Oana Alexandra Mocian and Ștefan Sorohan
J. Compos. Sci. 2024, 8(10), 426; https://doi.org/10.3390/jcs8100426 - 15 Oct 2024
Viewed by 670
Abstract
This work describes, through experimental and numerical investigations, the mechanical behavior and energy absorption characteristics of 3D-printed sandwich panels with cellular cores subjected to low-velocity impact. Using fused deposition modeling techniques (FDM), three different sandwich panels, one with a regular hexagonal core and [...] Read more.
This work describes, through experimental and numerical investigations, the mechanical behavior and energy absorption characteristics of 3D-printed sandwich panels with cellular cores subjected to low-velocity impact. Using fused deposition modeling techniques (FDM), three different sandwich panels, one with a regular hexagonal core and two with re-entrant cores at 0 and 90 degrees, were fabricated. The sandwich panels were subjected to low-velocity impact, at impact energies of 10 J and 15 J. A comprehensive investigation of the panels’ behavior through experimental testing and numerical simulation was conducted. The results indicate that the sandwich panel with a 90 degrees re-entrant core is stiffer and absorbs the largest amount of impact energy but, at the same time, suffers significant damage to the upper facesheet. The 0 degrees re-entrant core is compliant and provides both impact resistance and good energy absorption characteristics. Such a sandwich panel finds its application in the construction of personal protective equipment, where the aim is to minimize the forces transmitted during low-velocity impacts and maximize the total absorbed energy. Re-entrant core sandwich panels prove to be very good candidates for replacing the honeycomb core sandwich, depending on the desired engineering application. Full article
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<p>Geometrical configuration of the unit cells: (<b>a</b>) hexagonal and (<b>b</b>) re-entrant.</p>
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<p>Configuration of sandwich panels with (<b>a</b>) PS1 (0 degree re-entrant), (<b>b</b>) PS2 (90 degree re-entrant) and (<b>c</b>) PS3 (regular hexagonal) cores.</p>
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<p>Tensile specimens: (<b>a</b>) dimensions in mm and (<b>b</b>) tested specimens printed with a layer thickness of 0.2 mm.</p>
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<p>Average tensile engineering stress–strain curves for specimens printed with different layer thicknesses.</p>
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<p>Traction DIC analysis: (<b>a</b>) 0.15 mm, 0.2 mm and 0.3 mm layer thickness tested specimens and (<b>b</b>) principal strain at failure for 0.2 mm layer thickness.</p>
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<p>Average compression force-displacement curves for specimens printed with different layer thicknesses.</p>
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<p>Compression DIC test for a specimen with 0.2 mm layer thickness: (<b>a</b>) positioning of the specimen as observed with the two cameras and (<b>b</b>) principal strain variation at the end of the test.</p>
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<p>Measurement setup of the force pressing the sandwich panel: (<b>a</b>) load fixture and (<b>b</b>) measuring system of the force.</p>
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<p>Finite element model of the impact testing setup: 1—impactor, 2—clamping plate, 3—sandwich panel and 4—support.</p>
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<p>Force–time responses of sandwich panels under 10 J impact loading.</p>
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<p>Indentation of PLA sandwich panels after 10 J impact tests.</p>
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<p>Force–displacement responses of sandwich panels under 10 J impact loading.</p>
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<p>Energy–time histories of sandwich panels under 10 J Impact loading.</p>
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<p>Force–time responses of sandwich panels under 15 J impact loading.</p>
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<p>Force–displacement responses of sandwich panels under 15 J impact loading.</p>
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<p>Energy–time histories of sandwich panels under 15 J Impact loading.</p>
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<p>PLA sandwich panels after impact tests at 15 J.</p>
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<p>Comparison between experimental results and numerical analysis for panel PS1_10.</p>
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<p>Comparison between experimental results and numerical analysis for panel PS1_15.</p>
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<p>The equivalent von Mises stresses for the panel PS1_10.</p>
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<p>The equivalent von Mises stresses for the panel PS1_15.</p>
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<p>Comparison between experimental results and numerical analysis for panel PS2_10.</p>
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<p>Comparison between experimental results and numerical analysis for panel PS2_15.</p>
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<p>The equivalent von Mises stresses for the panel PS2_15.</p>
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<p>Comparison between experimental results and numerical analysis for panel PS3_10.</p>
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<p>Comparison between experimental results and numerical analysis for panel PS3_15.</p>
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<p>The equivalent von Mises stresses for the panel PS3_15.</p>
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13 pages, 7160 KiB  
Article
Experimental Analysis of the Mechanical Behavior of Shear Connectors for Precast Sandwich Wall Panels When Subjected to the Push-Out Tests
by John Kennedy Fonsêca Silva and Rodrigo de Melo Lameiras
Buildings 2024, 14(10), 3246; https://doi.org/10.3390/buildings14103246 - 14 Oct 2024
Viewed by 593
Abstract
Precast concrete sandwich panels consist of two outer layers connected by a central connector and an inner insulating layer that enhances thermal and acoustic performance. A key challenge with these panels is eliminating thermal bridges caused by metallic connectors, which reduce energy efficiency. [...] Read more.
Precast concrete sandwich panels consist of two outer layers connected by a central connector and an inner insulating layer that enhances thermal and acoustic performance. A key challenge with these panels is eliminating thermal bridges caused by metallic connectors, which reduce energy efficiency. PERFOFRP connectors, made from perforated glass fiber-reinforced polymer (GFRP) sheets, have been proposed to address this issue. These connectors feature holes that allow concrete to pass through, creating anchoring pins that enhance shear resistance and prevent the separation of the concrete layers. This research aimed to evaluate the effect of the diameter and number of holes on the mechanical strength of PERFOFRP connectors. Three diameters not previously reported in the literature were selected: 12.70 mm, 15.88 mm, and 19.05 mm. A total of 18 specimens, encompassing 6 different configurations with varying numbers of holes, underwent push-out tests. The most significant resistance increase was a 15% gain over non-perforated connectors, observed in the configuration featuring three holes of 19.05 mm. The connections exhibited rigid and nearly linear behavior until failure. Full article
(This article belongs to the Section Building Structures)
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<p>Shear strength mechanisms: (<b>a</b>) concrete front; (<b>b</b>) concrete dowel; (<b>c</b>) friction between the FRP and the concrete.</p>
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<p>Modes of failure of the PERFOFRP connector: (<b>a</b>) at the concrete front; (<b>b</b>) in the exposed FRP; (<b>c</b>) in the embedded FRP; (<b>d</b>) in the concrete dowels.</p>
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<p>Assembly of the specimens: (<b>a</b>) formworks; (<b>b</b>) concreted specimens; (<b>c</b>) specimen after removal of the insulating material; (<b>d</b>) specimen after fixing the steel plate to the connector.</p>
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<p>Geometry of the specimens, with the measurements presented in millimeters [mm]: (<b>a</b>) top view (sectional drawing), with the projection of the connector; (<b>b</b>) front view (sectional drawing), with the projection of the connector, highlighting the regions on the GFRP connector with (blue) and without steel reinforcement (yellow).</p>
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<p>Hole and spacing configuration of connectors, measured in millimeters [mm]: (<b>a</b>) SP-1-19.05; (<b>b</b>) SP-2-19.05; (<b>c</b>) SP-3-19.05; (<b>d</b>) SP-3-15.88; (<b>e</b>) SP-3-12.70; (<b>f</b>) SP-CTL.</p>
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<p>Push-out tests.</p>
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<p>Cracking patterns (the arrows, in the images where they appear, represent the top of the connector): (<b>a</b>) SP-1-19.05; (<b>b</b>) SP-2-19.05; (<b>c</b>) SP-3-19.05; (<b>d</b>) SP-3-15.88; (<b>e</b>) SP-3-12.70; (<b>f</b>) SP-CTL.</p>
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<p>Load versus relative displacement response of the specimens: (<b>a</b>) SP-12.70-1.75; (<b>b</b>) SP 12.70 2.00; (<b>c</b>) SP-12.70-2.50; (<b>d</b>) SP-12.70-3.00; (<b>e</b>) SP-CTL.</p>
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<p>Experimental results: for the ultimate load comparing (<b>a</b>) the number of holes and (<b>b</b>) the hole diameter; for the relative displacement comparing (<b>c</b>) the number of holes and (<b>d</b>) the hole diameter; and for the initial stiffness comparing (<b>e</b>) the number of holes and (<b>f</b>) the hole diameter.</p>
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18 pages, 32054 KiB  
Article
Study on the Process of Preparing Aluminum Foam Sandwich Panel Precursor by Friction Stir Welding
by Yu Zhang and Qiu Pang
Materials 2024, 17(20), 4981; https://doi.org/10.3390/ma17204981 - 11 Oct 2024
Viewed by 539
Abstract
In recent years, high-performance lightweight and multifunctional aluminum foam sandwiches (AFSs) can be successfully applied to spacecraft, automobiles, and high-speed trains. Friction stir welding (FSW) has been proposed as a new method for the preparation of AFS precursors in order to improve the [...] Read more.
In recent years, high-performance lightweight and multifunctional aluminum foam sandwiches (AFSs) can be successfully applied to spacecraft, automobiles, and high-speed trains. Friction stir welding (FSW) has been proposed as a new method for the preparation of AFS precursors in order to improve the cost-effectiveness and productivity of the preparation of AFS. In this study, the AFS precursors were prepared using the FSW process. The distribution of foaming agents in the AFS precursors and the structure and morphology of AFS were observed using optical microscopy (OM), scanning electron microscopy (SEM), and X-ray energy dispersive spectroscopy (EDS). The effects of the temperature and material flow on the distribution of the foaming agent during the FSW process were analyzed through experimental study and numerical simulation using ANSYS Fluent 19.0 software. The results show that the uniform distribution of the foaming agent in the matrix and excellent densification of AFS precursor can be prepared when the rotation speed is 1500 r/min, the travel speed is 25 mm/min, the tool plunge depth is 0.2 mm, and the tool moves along the retreating side (RS). In addition, the experimental and numerical simulations show that increasing the welding temperature improves the uniformity of foaming agent distribution and the area of AFS precursor prepared by single welding, shortening the thread length inhibits the foaming agent from reaching the upper sandwich plate and moving along the RS leads to a more uniform distribution of the foaming agent. Finally, the AFS with porosity of 74.55%, roundness of 0.97, and average pore diameter of 1.192 mm is prepared. Full article
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<p>Schematic diagram of the process of preparing AFS by FSW.</p>
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<p>Geometry, boundary definition, model mesh of the numerical model.</p>
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<p>Comparison of the metallographic organization of AFS precursors at the travel speed of 25 mm/min and different rotation speeds: (<b>a</b>) 1000 r/min, (<b>b</b>) 1500 r/min, (<b>c</b>) 2000 r/min.</p>
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<p>Comparison of metallographic organization of AFS precursor at the rotation speed of 25 mm/min and different travel speeds: (<b>a</b>) 25 mm/min, (<b>b</b>) 50 mm/min, (<b>c</b>) 100 mm/min.</p>
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<p>Comparison of metallographic organization of AFS precursors at different tool plunge depths: (<b>a</b>) 0.05 mm, (<b>b</b>) 0.20 mm, (<b>c</b>) 0.40 mm.</p>
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<p>The schematic diagram of tool movement along different directions and comparison of metallography: (<b>a</b>) the metallography of AS, (<b>b</b>) the metallography of RS, (<b>c</b>) the schematic diagram of AS, (<b>d</b>) the schematic diagram of RS.</p>
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<p>The formability experiment of AFS precursor under different welding tracks: (<b>a</b>) true stress-strain curves at different welding track, (<b>b</b>) Erikson formability test.</p>
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<p>Comparison of metallographic organization of AFS precursors at different pin thread lengths: (<b>a</b>) 4.00 mm, (<b>b</b>) 4.85 mm.</p>
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<p>Comparison of experimentally measured and numerically calculated temperature at 1500 r/min rotational speed and 25 mm/min travel speed: (<b>a</b>) experimentally measured schematic, (<b>b</b>) experimentally measured temperature curve, (<b>c</b>) numerically calculated temperature.</p>
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<p>Numerical simulation of the temperature distribution on the top surface of the workpiece for different parameters: (<b>a</b>) 25 mm/min travel speed and 1000 r/min rotational speed, (<b>b</b>) 100 mm/min travel speed and 1500 r/min rotational speed, (<b>c</b>) 25 mm/min travel speed and 1500 r/min rotational speed, (<b>d</b>) 25 mm/min travel speed and 2000 r/min rotational speed.</p>
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<p>The SEM morphology and EDS analysis of precursors at rotational speed 1500 r/min and tool movement along RS: (<b>a</b>) the metallographic organization of travel speed 100 mm/min, (<b>b</b>) SEM image of the rectangular region in (<b>a</b>), (<b>c</b>,<b>d</b>) the EDS maps corresponding to (<b>b</b>), (<b>e</b>) the metallographic organization of travel speed 25 mm/min, (<b>f</b>) SEM image of the rectangular region in (<b>e</b>), (<b>g</b>,<b>h</b>) the EDS maps corresponding to (<b>f</b>).</p>
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<p>Material flow in cross-section at different locations around the pin: (<b>a</b>) the cross-section schematic at different locations around the pin, (<b>b</b>) the metallographic organization of WZ, (<b>c</b>) the weld metal flow of cross-section I, (<b>d</b>) the weld metal flow of cross-section II, (<b>e</b>) the weld metal flow of cross-section III. The red arrows represent the material flow caused by the pin, the blue arrows represent the material flow caused by the tool shoulder, the yellow arrows represent the material flow caused by the material buildup, and the purple arrow represents material flowing out of the shoulder to form weld flash.</p>
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<p>The SEM morphology and EDS analysis of AFS precursor welded along the AS at travel speed 25 mm/min and rotational speed 1500 r/min: (<b>a</b>) metallographic organization, (<b>b</b>) SEM image of the rectangular region in (<b>a</b>), (<b>c</b>,<b>d</b>) the EDS maps corresponding to (<b>b</b>).</p>
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<p>SEM morphology and EDS analysis of AFS precursor under optimal welding parameters and methods (<b>a</b>) metallic organization of the precursor; (<b>b</b>) SEM image of the rectangular region in (<b>a</b>); (<b>c</b>) enlarged image of the rectangular region in (<b>b</b>); (<b>d</b>) EDS spectra of region A in (<b>c</b>); (<b>e</b>) EDS spectra of region B in (<b>c</b>); (<b>f</b>) EDS spectra of region C in (<b>c</b>); (<b>h</b>–<b>j</b>) EDS maps corresponding to (<b>c</b>).</p>
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<p>The foam microstructure of AFS precursors was prepared with pins of different thread lengths at 680 °C for different foaming times: (<b>a</b>,<b>b</b>) 4.00 mm (<b>c</b>,<b>d</b>) 4.85 mm.</p>
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<p>The bubble hole statistics and data comparison: (<b>a</b>) schematic of the bubble being counted, (<b>b</b>) comparison of bubble porosity, circularity, and diameter. Reprinted with permission from ref. [<a href="#B44-materials-17-04981" class="html-bibr">44</a>]. Copyright 2014 Elsevier. Reprinted with permission from ref. [<a href="#B45-materials-17-04981" class="html-bibr">45</a>]. Copyright 2018 J-Stage.</p>
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15 pages, 6728 KiB  
Article
Flexural Analysis of Additively Manufactured Continuous Fiber-Reinforced Honeycomb Sandwich Structures
by Rafael Guerra Silva, Esteban Gonzalez, Andres Inostroza and Gustavo Morales Pavez
J. Manuf. Mater. Process. 2024, 8(5), 226; https://doi.org/10.3390/jmmp8050226 - 10 Oct 2024
Viewed by 905
Abstract
This study explores the flexural behavior of continuous fiber-reinforced composite sandwich structures built entirely using material extrusion additive manufacturing. The continuous fiber additive manufacturing system used in this study works sequentially, thus enabling the addition of fiber reinforcement just in the face sheets, [...] Read more.
This study explores the flexural behavior of continuous fiber-reinforced composite sandwich structures built entirely using material extrusion additive manufacturing. The continuous fiber additive manufacturing system used in this study works sequentially, thus enabling the addition of fiber reinforcement just in the face sheets, where it is most effective. Three-point bending tests were carried out on sandwich panel specimens built using thermoplastic reinforced with continuous glass fiber to quantify the effect of fiber reinforcement and infill density in the flexural properties and failure mode. Sandwich structures containing continuous fiber reinforcement had higher flexural strength and rigidity than unreinforced sandwiches. On the other hand, an increase in the lattice core density did not improve the flexural strength and rigidity. The elastic modulus of fiber-reinforced 3D-printed sandwich panels exceeded the predictions of the analytical models; the equivalent homogeneous model had the best performance, with a 15% relative error. However, analytical models could not correctly predict the failure mode: wrinkle failure occurs at 75% and 30% of the critical load in fiber-reinforced sandwiches with low- and high-density cores, respectively. Furthermore, no model is currently available to predict interlayer debonding between the matrix and the thermoplastic coating of fiber layers. Divergences between analytical models and experimental results could be attributed to the simplifications in the models that do not consider defects inherent to additive manufacturing, such as air gaps and poor interlaminar bonding. Full article
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<p>Schematic representation of the dimensions of (<b>a</b>) specimen; (<b>b</b>) face sheet and core (including layer arrangement in sheet); and (<b>c</b>) hexagonal honeycomb.</p>
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<p>Universal testing machine model TIME WDW-200E with testing specimen.</p>
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<p>Comparison of force–displacement curves for different specimens for reinforced and non-reinforced sandwich structures.</p>
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<p>Average and range of (<b>a</b>) maximum load in bending (P<sub>max</sub>); (<b>b</b>) flexural modulus (E<sub>B</sub>); (<b>c</b>) facing stress (σ), as in Equation (2).</p>
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<p>Average maximum load before failure in bending (P<sub>max</sub>) and the failure load under face yielding and face wrinkling for reinforced and non-reinforced specimens.</p>
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<p>Damage in low-density reinforced specimens (R4i25) (<b>a</b>) Fiber breakage in the bottom facing (R4i25-A) (<b>b</b>) Crack across the core (R4i25-A) (<b>c</b>) Wrinkling in the top face (R4i25-B) (<b>d</b>) Delamination between the face and core in the top side (R4i25-C).</p>
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<p>Damage in high-density reinforced specimens (R4i50) (<b>a</b>) Crack on the lower face (R4i50-B) (<b>b</b>) Crack on the lower face and debonding on the upper face (R4i50-B) (<b>c</b>) Wrinkling on the upper face (R4i50-C) (<b>d</b>) Debonding on the upper face (R4i50-A).</p>
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26 pages, 29595 KiB  
Article
Induction Heating of Laminated Composite Structures with Magnetically Responsive Nanocomposite Interlayers for Debonding-on-Demand Applications
by Eleni Gkartzou, Konstantinos Zafeiris, Christos Tsirogiannis, Alberto Pedreira, Adrián Rodríguez, Pablo Romero-Rodriguez, Giorgos P. Gakis, Tatjana Kosanovic-Milickovic, Apostolos Kyritsis and Costas A. Charitidis
Polymers 2024, 16(19), 2760; https://doi.org/10.3390/polym16192760 - 30 Sep 2024
Viewed by 1027
Abstract
In the present study, the feasibility to achieve localized induction heating and debonding of multi-material composite structures is assessed in testing coupons prepared by Automated Fiber Placement (AFP) and extrusion-based additive manufacturing (AM) technologies. Nano-compounds of Polyether-ketone-ketone (PEKK) with iron oxide nanoparticles acting [...] Read more.
In the present study, the feasibility to achieve localized induction heating and debonding of multi-material composite structures is assessed in testing coupons prepared by Automated Fiber Placement (AFP) and extrusion-based additive manufacturing (AM) technologies. Nano-compounds of Polyether-ketone-ketone (PEKK) with iron oxide nanoparticles acting as electromagnetic susceptors have been processed in a parallel co-rotating twin-screw extruder to produce filament feedstock for extrusion-based AM. The integration of nanocomposite interlayers as discrete debonding zones (DZ) by AFP-AM manufacturing has been investigated for two types of sandwich-structured laminate composites, i.e., laminate-DZ-laminate panels (Type I) and laminate-DZ-AM gyroid structures (Type II). Specimens were exposed to an alternating magnetic field generated by a radio frequency generator and a flat spiral copper induction coil, and induction heating parameters (frequency, power, heating time, sample standoff distance from coil) have been investigated in correlation with real-time thermal imaging to define the debonding process window without compromising laminate quality. For the optimized process parameters, i.e., 2–3 kW generator power and 20–25 mm standoff distance, corresponding to magnetic field intensities in the range of 3–5 kA m−1, specimens were effectively heated above PEKK melting temperature, exhibiting high heating rates within the range of 5.3–9.4 °C/s (Type I) and 8.0–17.5 °C/s (Type II). The results demonstrated that localized induction heating successfully facilitated debonding, leading to full unzipping of the debonding zones in both laminate structures. Further insight on PEKK nanocomposites debonding performance was provided by thermal, morphological characterization and non-destructive inspection via X-ray micro-computed tomography at different processing stages. The developed framework aims to contribute to the development of rapid, on-demand joining, repair and disassembly technologies for thermoplastic composites, towards more efficient maintenance, repair and overhaul operations in the aviation sector and beyond. Full article
(This article belongs to the Special Issue Polymeric Materials and Their Application in 3D Printing, 2nd Edition)
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Graphical abstract
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<p>Schematic of the mounting setup for specimen static heating and debonding trials.</p>
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<p>Induction heating testing setup.</p>
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<p>Simulation results for magnetic flux density norm (<b>a</b>,<b>b</b>) and magnetic field intensity (<b>c</b>,<b>d</b>) for 30 mm standoff distance (top/side coil view).</p>
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<p>Indicative images of agglomerate distribution for PEKK + Fe<sub>3</sub>O<sub>4</sub> 7.5 wt% FFF filament sample: (<b>a</b>) XZ, XY, ZY cross sections of reconstructed grayscale slices (scale bar: 800 μm); (<b>b</b>,<b>c</b>) 3D visualization of sample volume (scale bar: 250 μm, 2.5 μm voxel size).</p>
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<p>(<b>a</b>) Color-coded 2D visualization of agglomerate size (scale = 300 μm); (<b>b</b>) agglomerate size classification employing the volume-equivalent sphere diameter model (PEKK + Fe<sub>3</sub>O<sub>4</sub> 7.5 wt%); (<b>c</b>) distribution of agglomerate separation distances for PEKK + Fe<sub>3</sub>O<sub>4</sub> 7.5 wt%.</p>
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<p>Comparative TGA thermograms for each processing stage, namely pellet form (S1), FFF interlayer (S2), after re-melting with induction heating (S3).</p>
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<p>DSC thermograms for each processing stage, namely pellet form (S1), FFF interlayer (S2), after re-melting with induction heating (S3). (<b>a</b>) Cooling cycle; (<b>b</b>) 2nd scan heating cycles.</p>
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<p>Three-dimensional visualization of Type I (<b>a</b>,<b>b</b>) and Type II (<b>c</b>,<b>d</b>) specimens and FFF nanocomposite interlayer before induction heating (white regions: FFF interlayer and NP agglomerates, gray region: CFRP/polymer matrix, black: background/air; scale bar: 1 mm).</p>
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<p>Absolute value of the induced magnetic field intensity (H) simulated for different standoff distances (<b>a</b>) and generator power/frequency (<b>b</b>) settings; (<b>c</b>–<b>e</b>) magnetic field intensity simulated for the FFF debonding zone located at different standoff distances from the coil.</p>
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<p>(<b>a</b>) Top-view image of the automated movement setup captured with the IR camera; (<b>b</b>) measurement of reference PAEK sample without nanocomposite FFF interlayer—no increase in specimen temperature recorded.</p>
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<p>Representative induction heating curves of Type I specimens at 2 kW power value and varying standoff distance values (20–45 mm).</p>
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<p>Representative induction heating curves of Type I specimens at 3 kW power value and varying standoff distance values (20–45 mm).</p>
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<p>Representative induction heating curves of Type I specimens at 20 mm standoff distance and varying power values (2 kW–3 kW).</p>
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<p>Representative induction heating curves of Type II (gyroid) specimens at 20 mm standoff distance and varying power values (2 kW–3 kW).</p>
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<p>Indicative thermal camera images of debonding of Type I (<b>a</b>) and Type II (<b>b</b>) specimens, coupled with specimen structure prior and after debonding.</p>
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<p>(<b>a</b>–<b>e</b>) Cross section images of debonded CFRP laminates extracted from Type I specimens processed under 2–3 kW generator power range at 20 mm standoff distance (scale bar: 750 μm).</p>
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<p>Top-view images of debonded surface morphology for 2 kW (<b>a</b>) and 3 kW (<b>b</b>) generator power at 20 mm standoff distance (scale bar: 1 mm).</p>
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<p>Top (<b>a</b>) and side (<b>b</b>) view of debonded gyroid structure and CFRP laminate (<b>c</b>) of Type II specimen processed under 2 kW generator power at 20 mm standoff distance (scale bar: 2.5 mm).</p>
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<p>SEM morphology images (<b>a</b>–<b>d</b>) and EDS elemental analysis (<b>e</b>,<b>f</b>) of the intermediate layer (debonding zone) of debonded specimen.</p>
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<p>Three-dimensional visualization of CFRP laminates (<b>a</b> and <b>b</b>, <b>c</b> and <b>d</b>) and residual FFF nanocomposite interlayer after induction heating and debonding (white regions: NP agglomerates, gray region: CFRP/polymer matrix, black: background/air). Top (<b>a</b>,<b>c</b>)/bottom (<b>b</b>,<b>d</b>) images depict the same samples, with adjustment of the attenuation coefficient range to isolate features with different absorptivity (scale bar: 1 mm).</p>
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<p>(<b>a</b>) Simulated geometry for the induction heating model; (<b>b</b>) computational mesh used to discretize the computational domain.</p>
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21 pages, 35079 KiB  
Article
Energy Absorption Properties of 3D-Printed Polymeric Gyroid Structures for an Aircraft Wing Leading Edge
by Mats Overbeck, Sebastian Heimbs, Jan Kube and Christian Hühne
Aerospace 2024, 11(10), 801; https://doi.org/10.3390/aerospace11100801 - 29 Sep 2024
Viewed by 1107
Abstract
Laminar flow offers significant potential for increasing the energy efficiency of future transport aircraft. At the Cluster of Excellence SE2A—Sustainable and Energy-Efficient Aviation—the laminarization of the wing by means of hybrid laminar flow control (HLFC) is being investigated. The aim is [...] Read more.
Laminar flow offers significant potential for increasing the energy efficiency of future transport aircraft. At the Cluster of Excellence SE2A—Sustainable and Energy-Efficient Aviation—the laminarization of the wing by means of hybrid laminar flow control (HLFC) is being investigated. The aim is to maintain the boundary layer as laminar for up to 80% of the chord length of the wing. This is achieved by active suction on the leading edge and the rear part of the wing. The suction panels are constructed with a thin micro-perforated skin and a supporting open-cellular core structure. The mechanical requirements for this kind of sandwich structure vary depending on its position of usage. The suction panel on the leading edge must be able to sustain bird strikes, while the suction panel on the rear part must sustain bending loads from the deformation of the wing. The objective of this study was to investigate the energy absorption properties of a triply periodic minimal surface (TPMS) structure that can be used as a bird strike-resistant core in the wing leading edge. To this end, cubic-sheet-based gyroid specimens of different polymeric materials and different geometric dimensions were manufactured using additive manufacturing processes. The specimens were then tested under quasi-static compression and dynamic crushing loading until failure. It was found that the mechanical behavior was dependent on the material, the unit cell size, the relative density, and the loading rate. In general, the weight-specific energy absorption (SEA) at 50% compaction increased with increasing relative density. Polyurethane specimens exhibited an increase in SEA with increasing loading rate, as opposed to the specimens of the other investigated polymers. A smaller unit cell size induced a more consistent energy absorption, due to the higher plateau force. Full article
(This article belongs to the Special Issue Advanced Aerospace Composite Materials and Smart Structures)
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<p>Schematic illustration of the gyroid structure on the wing leading edge with front spar (<b>a</b>), TPMS-based sandwich core structure (<b>b</b>), and micro-perforated skin (<b>c</b>). The air flow (<b>d</b>) over the micro-perforated skin draws the boundary layer into the leading edge (<b>e</b>). The orientation in flight direction poses a risk of bird strike (<b>f</b>).</p>
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<p><span class="html-italic">Polyurethane Rigid 650</span> gyroid structures with a relative density of 23% and different unit cell sizes ranging from 7 mm to 20 mm.</p>
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<p><span class="html-italic">Nylon 11</span> gyroid structures with a unit cell size of 20 mm and different relative densities ranging from 7% to 23%.</p>
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<p>Test setup for quasi-static (<b>a</b>) and dynamic (<b>b</b>) loading.</p>
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<p>Force–displacement curves of <span class="html-italic">Standard Resin</span> gyroid specimens with different relative densities <math display="inline"><semantics> <msup> <mi>ρ</mi> <mo>∗</mo> </msup> </semantics></math> ranging from 7% to 23% under quasi-static compression and dynamic crushing loading.</p>
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<p>Weight-specific energy absorption over displacement of <span class="html-italic">Standard Resin</span> gyroid specimens with different relative densities <math display="inline"><semantics> <msup> <mi>ρ</mi> <mo>∗</mo> </msup> </semantics></math> ranging from 7% to 23% under quasi-static compression and dynamic crushing loading.</p>
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<p>Zoomed in force–displacement (<b>a</b>,<b>b</b>) and SEA–displacement (<b>c</b>,<b>d</b>) curves of <span class="html-italic">Standard Resin</span> gyroid specimens with relative densities <math display="inline"><semantics> <msup> <mi>ρ</mi> <mo>∗</mo> </msup> </semantics></math> of 7% (<b>a</b>,<b>c</b>) and 12% (<b>b</b>,<b>d</b>) under quasi-static compression and dynamic crushing loading.</p>
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<p>Force–displacement (<b>a</b>) and weight-specific energy absorption (<b>b</b>) curves of <span class="html-italic">Polyurethane Rigid 650</span> gyroid specimens with a relative density <math display="inline"><semantics> <msup> <mi>ρ</mi> <mo>∗</mo> </msup> </semantics></math> of 23% under quasi-static compression and dynamic crushing loading.</p>
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<p>Failure behavior of <span class="html-italic">Polyurethane Rigid 650</span> gyroid specimens of 12% relative density <math display="inline"><semantics> <msup> <mi>ρ</mi> <mo>∗</mo> </msup> </semantics></math> with a unit cell size of 10<math display="inline"><semantics> <mi mathvariant="normal">m</mi> </semantics></math><math display="inline"><semantics> <mi mathvariant="normal">m</mi> </semantics></math> under quasi-static compression (<b>a</b>) and dynamic crushing (<b>b</b>) loading at a compaction of 50%.</p>
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<p>Force–displacement curves of <span class="html-italic">Nylon 11</span> gyroid specimens with different relative densities <math display="inline"><semantics> <msup> <mi>ρ</mi> <mo>∗</mo> </msup> </semantics></math> ranging from 7% to 23% under quasi-static compression and dynamic crushing loading.</p>
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<p>Weight-specific energy absorption over displacement of <span class="html-italic">Nylon 11</span> gyroid specimens with different relative densities <math display="inline"><semantics> <msup> <mi>ρ</mi> <mo>∗</mo> </msup> </semantics></math> ranging from 7% to 23% under quasi-static compression and dynamic crushing loading.</p>
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<p>Failure behavior of <span class="html-italic">Nylon 11</span> gyroid specimens of 12% relative density with a unit cell size of 20 <math display="inline"><semantics> <mi mathvariant="normal">m</mi> </semantics></math><math display="inline"><semantics> <mi mathvariant="normal">m</mi> </semantics></math> under quasi-static compression (<b>a</b>,<b>b</b>) and dynamic crushing (<b>c</b>,<b>d</b>) loading.</p>
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<p>Plateau force of <span class="html-italic">Standard Resin</span> (<b>a</b>) <span class="html-italic">Polyurethane Rigid 650</span> (<b>b</b>), and <span class="html-italic">Nylon 11</span> (<b>c</b>) gyroid specimens over the relative density with constant unit cell size.</p>
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<p>Weight-specific energy absorption at a compaction of 50% of gyroid specimens over the relative density.</p>
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19 pages, 9739 KiB  
Article
Lateral Performance of Composite Wall with Cold-Formed Thin-Walled Steel–Concrete Sandwich Panel
by Jian Zou, Baozhu Cao, Xiang Zeng and Yuchuan Zhang
Buildings 2024, 14(9), 2928; https://doi.org/10.3390/buildings14092928 - 16 Sep 2024
Viewed by 715
Abstract
To study the lateral performance of a cold-formed steel–concrete insulation sandwich panel composite wall, two full-scale specimens with different arrangements were designed. The specimens underwent cyclic loading tests to examine the failure characteristics of the composite wall, and lateral performance aspects such as [...] Read more.
To study the lateral performance of a cold-formed steel–concrete insulation sandwich panel composite wall, two full-scale specimens with different arrangements were designed. The specimens underwent cyclic loading tests to examine the failure characteristics of the composite wall, and lateral performance aspects such as the experimental hysteresis curve, skeleton curve, and characteristic value of the whole loading process were acquired. The experimental results indicate that the failure of the composite wall system was primarily caused by the failure of the connection; the overall lateral performance of composite walls with one wall panel at the bottom and two wall panels at the top (W1) was superior to that of composite walls with two wall panels at the bottom and one wall panel at the top (W2). When loaded to an inter-story drift ratio of 1/300, the composite wall did not exhibit any significant damage. A finite element (FE) model was developed and validated by the experiments. Factors affecting the shear bearing capacity were analyzed based on the FE model, including the yield strength of diagonal braces, the thickness of the diagonal braces, the arrangement pattern of the wall panels, the dimensions of the wall panels, and the strength of the connection of the L-shaped connector and the flat connector. The FE results show that all these factors can influence the lateral performance of the composite wall. Full article
(This article belongs to the Section Building Structures)
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<p>Composite wall: (<b>a</b>) flat steel connectors; (<b>b</b>) axial view of L-shaped connectors; (<b>c</b>) W1; (<b>d</b>) W2.</p>
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<p>Joint and structure of upper and lower wall panels.</p>
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<p>Mechanical property testing of the material: (<b>a</b>) steel tensile fracture; (<b>b</b>) concrete compressive failure; (<b>c</b>) shear failure of self-tapping screws.</p>
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<p>Loading regime.</p>
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<p>Layout of the loading apparatus and sensor position: (<b>a</b>) loading apparatus schematic; (<b>b</b>) sensor position layout diagram.</p>
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<p>Failure patterns of specimen W1: (<b>a</b>) the upper and lower wall panels are out of alignment; (<b>b</b>) flat connector at the left end is damaged; (<b>c</b>) the upper wall panels are detached. (<b>d</b>) The wall panel bulged outward.</p>
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<p>Failure patterns of specimen W2: (<b>a</b>) the L-shaped connector at the right end is damaged; (<b>b</b>) the upper and lower wall panels are out of alignment; (<b>c</b>) flat connector at the left end is damaged.</p>
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<p>Hysteresis curves of the specimens: (<b>a</b>) specimen W1 hysteresis curve; (<b>b</b>) specimen W2 hysteresis curve.</p>
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<p>Skeleton curve.</p>
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<p>Stiffness degradation curves.</p>
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<p>Computation of energy dissipation coefficient.</p>
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<p>Load–displacement curve of steel and concrete connection.</p>
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<p>Finite element modeling: (<b>a</b>) finite element global model; (<b>b</b>) finite element model of steel framework.</p>
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<p>The finite element failure mode of the W1 specimen: (<b>a</b>) Von Mises stress contour map of steel frame; (<b>b</b>) Von Mises stress contour map of wall panel; (<b>c</b>) displacement contour map in the Uy direction of steel frame; (<b>d</b>) displacement contour map in the Uy direction of wall panel.</p>
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<p>Comparison of experimental and finite element simulation results: (<b>a</b>) skeleton curve of specimen W1; (<b>b</b>) skeleton curve of specimen W2.</p>
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<p>The influence of diagonal brace yield strength on seismic performance: (<b>a</b>) skeletal curves with different diagonal brace yield strengths. (<b>b</b>) The effect on the shear capability of the specimens.</p>
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<p>The impact of diagonal bracing thickness on seismic performance. (<b>a</b>) Skeletal curves with varying diagonal brace thicknesses; (<b>b</b>) the impact on the load-bearing capacity of the specimens.</p>
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<p>Arrangement pattern of the wall panels. (<b>a</b>) W1-A1; (<b>b</b>) W1-A2; (<b>c</b>) W1-A3.</p>
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<p>The skeleton curves for different wall panel arrangements.</p>
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<p>Dimensions of wall panels. (<b>a</b>) W1-R-1. (<b>b</b>) W1-R-2. (<b>c</b>) W1-R-3.</p>
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<p>The skeleton curves for different dimensions of wall panels.</p>
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<p>The skeleton curves for different strengths of the connection.</p>
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24 pages, 40305 KiB  
Article
Numerical Investigation on Dynamic Response of Carbon Fiber Honeycomb Sandwich Panels Subject to Underwater Impact Load
by Cheng Zheng, Mingfei Wang, Yiwen Wang, Yawen Liao, Xiangshao Kong and Weiguo Wu
J. Mar. Sci. Eng. 2024, 12(9), 1513; https://doi.org/10.3390/jmse12091513 - 2 Sep 2024
Viewed by 896
Abstract
This study investigates the dynamic response and failure mechanisms of carbon fiber honeycomb sandwich structures under underwater impact loads using finite element numerical simulation. The geometric modeling was performed using HyperMesh, and the dynamic response simulations were carried out in ABAQUS, focusing on [...] Read more.
This study investigates the dynamic response and failure mechanisms of carbon fiber honeycomb sandwich structures under underwater impact loads using finite element numerical simulation. The geometric modeling was performed using HyperMesh, and the dynamic response simulations were carried out in ABAQUS, focusing on honeycomb core configurations with varying edge lengths, heights, and gradient forms. The Hashin damage model was employed to describe the damage evolution of the composite materials. The simulation results revealed that the dynamic response was significantly influenced by the initial shock wave pressure and the geometrical parameters of the honeycomb cells. Larger cell-edge lengths and heights generally resulted in improved energy absorption and reduced rear panel displacement. Among the different configurations, interlayer gradient honeycomb structures demonstrated superior impact resistance compared to homogeneous and in-plane gradient structures, particularly under higher initial shock wave pressures. These findings contribute to optimizing the design of carbon fiber honeycomb sandwich structures for enhanced impact resistance in relevant applications. Full article
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<p>Schematic diagram of finite element model of carbon fiber honeycomb sandwich structure and watershed structure.</p>
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<p>Schematic diagram of interlayer gradient honeycomb sandwich structure.</p>
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<p>Schematic diagram of in-plane gradient honeycomb sandwich structure.</p>
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<p>Peak shock wave pressure in the water and the change rule of carbon fiber rear panel center deformation with different mesh sizes. (<b>a</b>) Variation law of peak pressure of shock wave in water under different mesh sizes. (<b>b</b>) Deformation law of carbon fiber structural panels under different mesh sizes.</p>
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<p>Comparison of simulation and test results of the center displacement of the rear panel of the structure at 8 g and 15 g conditions.</p>
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<p>Damage (shear damage) diagram of carbon fiber honeycomb sandwich structure.</p>
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<p>Stress distribution of carbon fiber honeycomb sandwich structure at different times.</p>
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<p>Time history curve of displacement of rear panel of carbon fiber honeycomb sandwich structure.</p>
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<p>Total energy, percentage of energy absorbed by each component and damage absorption of the carbon fiber honeycomb sandwich structure for Case 4 (initial pressure 120 MPa, H = 15 mm, L = 20 mm, T = 2 mm). (<b>a</b>) Total energy of carbon fiber honeycomb sandwich structure and energy absorption ratio of each component. (<b>b</b>) Specific energy absorption of honeycomb structure under initial pressure of 120 MPa.</p>
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<p>Specific energy absorption of honeycomb structure and rear panel center displacement under different initial pressures and different edge lengths of honeycomb cells. (<b>a</b>) Specific energy absorption of honeycomb structure under different initial pressures. (<b>b</b>) Peak displacement of the rear panel center under different initial pressures.</p>
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<p>Specific energy absorption efficiency of honeycomb structure with different initial pressures and different edge lengths of honeycomb cells.</p>
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<p>Total energy, percentage of energy absorbed by each component, and damage absorption of carbon fiber honeycomb sandwich structure in Case VII. (<b>a</b>) Total energy of carbon fiber honeycomb sandwich structure and energy absorption ratio of each component. (<b>b</b>) Specific energy absorption of honeycomb structure under the initial pressure of 120 MPa.</p>
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<p>Specific energy absorption diagram of honeycomb structure and peak displacement of rear panel center at different initial pressures and different heights of honeycomb cells. (<b>a</b>) Specific energy absorption of honeycomb structure under different initial pressures. (<b>b</b>) Peak displacement of the rear panel center under different initial pressures.</p>
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<p>Specific energy absorption efficiency of honeycomb structure under different initial pressures and different heights of honeycomb cell elements.</p>
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<p>Gradient carbon fiber honeycomb sandwich structure total energy and the percentage of energy absorbed by each component. (<b>a</b>) Total energy of interlayer gradient carbon fiber honeycomb sandwich structure and energy absorption ratio of each component. (<b>b</b>) Total energy of in-plane gradient carbon fiber honeycomb sandwich structure and energy absorption ratio of each component.</p>
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<p>Variation of specific absorption energy with initial pressure of shock wave in water for different honeycomb sandwich structures.</p>
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<p>Variation of energy absorption percentage with initial pressure of shock wave in water for different honeycomb sandwich structures.</p>
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<p>Variation of rear panel center displacement with initial shock wave pressure in water for different honeycomb sandwich structures.</p>
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<p>Variation of specific energy absorption efficiency with initial pressure of shock wave in water for different honeycomb sandwich structures.</p>
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