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Search Results (1,401)

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25 pages, 1710 KiB  
Article
Local Stability Analysis of a Composite Corrugated Steel Plate Pipe-Arch in Soil
by Chengwen Che, Pingping Hu, Feng Shi, Pengsen Xu, Junxiu Liu and Kai Li
Buildings 2024, 14(10), 3290; https://doi.org/10.3390/buildings14103290 - 17 Oct 2024
Viewed by 206
Abstract
The straight part of the corrugated steel plate (CSP) pipe-arch structure in soil may cause local buckling instability due to insufficient load-bearing capacity. Recently, composite CSP pipe-arch has been widely utilized to enhance structural stability, and their stability needs to be thoroughly investigated. [...] Read more.
The straight part of the corrugated steel plate (CSP) pipe-arch structure in soil may cause local buckling instability due to insufficient load-bearing capacity. Recently, composite CSP pipe-arch has been widely utilized to enhance structural stability, and their stability needs to be thoroughly investigated. This paper studies the local buckling stability problem of the straight part of composite CSP pipe-arch in soil by simplifying the soil support and introducing the inter-layer bonding effect. Based on elastic stability theory, a theoretical mechanical model of composite CSP pipe-arch was proposed. The Rayleigh–Ritz method and the semi-combined composite structure stiffness approximation were used to derive the critical buckling conditions for the straight part of the composite CSP pipe-arch. Through numerical calculation and influencing factors analysis, it is concluded that the critical buckling load of the straight part of the composite CSP pipe-arch structure is affected by the elastic modulus, thickness, Poisson’s ratio, rotational restraint stiffness and side length of the straight part of the material. In particular, it is found that as the inter-layer bonding coefficient increases, the critical buckling load is improved, while the critical buckling wave number is mainly influenced by the width of the straight part, elastic modulus, and inter-layer bonding coefficient. Additionally, we discussed the coupling effect of several key parameters on the stability of the structure. The results of this study offer theoretical foundations and guidance for the application of composite CSP pipe-arch in soil engineering, such as culverts, tunnels, and pipeline transportation. Full article
(This article belongs to the Section Building Structures)
17 pages, 9762 KiB  
Article
Experimental and Numerical Analysis of Bolted Repair for Composite Laminates with Delamination Damage
by Shan Xiao, Mingxuan Huang, Zhonghai Xu, Yusong Yang and Shanyi Du
Polymers 2024, 16(20), 2918; https://doi.org/10.3390/polym16202918 - 17 Oct 2024
Viewed by 279
Abstract
Composite materials are widely used in aircraft due to the urgent need for high-quality structures in aerospace engineering. In order to verify the effectiveness of complex bolt repairs on composite structures, compression tests have been performed on three types (intact, damaged, and repaired) [...] Read more.
Composite materials are widely used in aircraft due to the urgent need for high-quality structures in aerospace engineering. In order to verify the effectiveness of complex bolt repairs on composite structures, compression tests have been performed on three types (intact, damaged, and repaired) of composite plate specimens, and finite element simulation results of these three types’ specimens were obtained. The experimental results show that for damaged composite laminates, the strength recovery after bolt repair can reach an impressive 107%, and the delamination propagation caused by over-buckling deformation is considered to be the main cause of failure, which also suggests that although bolt repair can improve the strength of the specimens, it has a limited ability to inhibit delamination propagation. The simulation results of the finite element model in this paper are in good agreement with the actual experimental results, and the maximum error does not exceed 7.9%. In conclusion, this paper verifies the suitability of the proposed repair scheme in engineering applications and the correctness of the modeling method for repaired composite laminates. Full article
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<p>Linear degradation model used in Abaqus.</p>
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<p>Composite laminates with intact (<b>a</b>) and pre-existing delamination damage (<b>b</b>) and repaired with bolts (<b>c</b>).</p>
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<p>Distribution situation and numbers of strain gauge: intact (<b>a</b>) and pre-existing delamination (<b>b</b>) and repaired with bolted joint (<b>c</b>).</p>
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<p>Repair operation flowchart.</p>
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<p>Experimental setup.</p>
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<p>Finite element model for initial composite laminates, where green is the laminate model and white is the fixture model.</p>
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<p>Finite element model for composite laminates with delamination.</p>
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<p>Finite element model of repaired delaminated composite laminates.</p>
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<p>Intact composite laminates after compression test ((<b>a</b>–<b>c</b>) correspond to specimens 1, 2, and 3, respectively).</p>
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<p>Comparison of experimental and simulation results.</p>
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<p>Damaged composite laminates before (<b>a</b>) and after (<b>b</b>) compression test.</p>
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<p>Images of the repaired composite laminates before (<b>a</b>) and after (<b>b</b>) compression test.</p>
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<p>Strain response of initial composite laminates.</p>
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<p>Strain response of damaged composite laminates.</p>
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<p>Strain response of repaired composite laminate.</p>
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<p>Fiber damage cloud map and experimental result graph.</p>
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<p>Predicted displacement contour map and experimental result.</p>
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<p>Calculation results of repaired compression test specimen.</p>
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15 pages, 3633 KiB  
Article
Rib Alignment Control of Long-Span Arch Bridge in Cable-Stayed Buckle by Multi-Objective Optimization
by Mengsheng Yu, Xinyu Yao, Longlin Wang, Tianzhi Hao and Nianchun Deng
Buildings 2024, 14(10), 3281; https://doi.org/10.3390/buildings14103281 - 17 Oct 2024
Viewed by 253
Abstract
The construction duration of long-span arch bridges is excessively prolonged due to insufficient closing precision and the non-convergence of traditional cable adjustment calculation methods. This study investigates cable force management in long-span concrete-filled steel tubular (CFST) arch bridges during cable-stayed buckle construction, aiming [...] Read more.
The construction duration of long-span arch bridges is excessively prolonged due to insufficient closing precision and the non-convergence of traditional cable adjustment calculation methods. This study investigates cable force management in long-span concrete-filled steel tubular (CFST) arch bridges during cable-stayed buckle construction, aiming to improve construction safety and precision in arch rib alignment. Using the Pingnan Third Bridge and Tian’e Longtan Bridge as practical examples, the research develops a multi-objective optimization method for cable forces that integrates influence matrices, constrained minimization, and a forward iterative approach. This method offers a robust strategy for tensioning and cable-stayed buckling, enabling real-time monitoring, calculation, and adjustment during the construction of large-span CFST arch bridges. The results reveal that the iterative approach notably enhances calculation efficiency compared to conventional methods. For instance, field measurements at the Pingnan Third Bridge show a minimal arch closure error of only 3 mm. Additionally, the study addresses concerns about excessive stress in exposed steel tubes during concrete casting. By optimizing the sequence of main arch closure and concrete casting, stress in the exposed steel tube is reduced from 373 MPa to 316 MPa, thus meeting specification requirements. In summary, the multi-objective cable force optimization method demonstrates superior efficiency in determining cable tension and controlling rib alignment during cable-stayed buckle construction of long-span CFST arch bridges. Full article
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<p>Iterative calculation steps of cable force adjustment.</p>
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<p>The elevation view of the Pingnan Third Bridge.</p>
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<p>The schematic diagram of the displacement control point.</p>
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<p>Elevation of Tian’e Longtan Grand Bridge.</p>
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<p>The schematic diagram of the cable-stayed buckle construction.</p>
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<p>Relationship curve between dispersion and delta of the Pingnan Third Bridge.</p>
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<p>Relationship curve between dispersion and delta of the Tian’e Longtan Bridge.</p>
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<p>Arch rib segment division (Pingnan Third Bridge) (the number indicates the arch rib segment number).</p>
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<p>Arch rib segment division (Tian’e Longtan Bridge) (the number indicates the arch rib segment number).</p>
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<p>The comparison diagram of the initial iteration cable tension and optimized cable force of Pingnan Third Bridge (unit: kN). Note: S = South, N = North, A = Cable force after upstream optimization, B = Cable force after downstream optimization, C = Cable force before upstream optimization, D = Cable force before downstream optimization.</p>
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<p>The comparison diagram of the initial iteration cable tension and optimized cable force of the Tian’e Longtan Bridge (unit: kN). Note: S = South, N = North, A = Cable force after upstream optimization, B = Cable force after downstream optimization, C = Cable force before upstream optimization, D = Cable force before downstream optimization.</p>
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<p>The displacement result diagram of the Pingnan Third Bridge. Note: U = Up, D = Down, A = The difference between the alignment of the after cable closure and the target displacement, B = The displacement difference between south and north after cable closure, C = The displacement difference between the upstream and downstream of the installed horizontal brace segment, D = The displacement difference between south and north in the current tension stage.</p>
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<p>The displacement result diagram of the Tian’e Longtan Bridge. Note: U = Up, D = Down, A = The difference between the alignment of the after cable closure and the target displacement, B = The displacement difference between south and north after cable closure, C = The displacement difference between south and north in the current tension stage, D = The displacement difference between the upstream and downstream of the installed horizontal transverse brace segment.</p>
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<p>The displacement of in situ measured and theoretical of the Pingnan Third Bridge. Note: S = South, N = North, A = The upstream in situ measured displacement, B = The upstream theoretical displacement, C = The downstream in situ measured displacement, D = The downstream theoretical displacement.</p>
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<p>The comparison diagram of stress optimization schemes for the Tian’e Longtan Bridge.</p>
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15 pages, 3339 KiB  
Article
Post-Buckling Response of Carbon/Epoxy Laminates with Delamination under Quasi-Static Compression: Experiments and Numerical Simulations
by Fei Xia, Zikun Wang, Yi Wang, Heqing Liu and Jianghong Xue
Materials 2024, 17(20), 5047; https://doi.org/10.3390/ma17205047 (registering DOI) - 15 Oct 2024
Viewed by 303
Abstract
Delamination is a common type of damage in composite laminates that can significantly affect the integrity and stability of structural components. This study investigates the post-buckling behavior of carbon fiber-reinforced epoxy composite laminates with embedded delamination under quasi-static compression. Experimental tests were conducted [...] Read more.
Delamination is a common type of damage in composite laminates that can significantly affect the integrity and stability of structural components. This study investigates the post-buckling behavior of carbon fiber-reinforced epoxy composite laminates with embedded delamination under quasi-static compression. Experimental tests were conducted using an electronic universal material testing machine to measure deformation and load-bearing capacity in the post-buckling stage. The specimens, prepared from T300 carbon fiber and TDE-85 epoxy resin prepreg, were subjected to axial compressive loads with delamination simulated by embedding Teflon films. Finite element analysis (FEA) was performed using ABAQUS software, incorporating a four-part model to simulate delaminated structures, with results validated against experimental data through comprehensive convergence analysis. The findings reveal that increasing delamination depth and length decrease overall stiffness, leading to an earlier onset of buckling. Structural instability was observed to vary with the size of delamination, while the post-buckling deformation mode consistently exhibited a half-wave pattern. This research underscores the critical impact of delamination on the structural integrity and load-bearing performance of composite laminates, providing essential insights for developing more effective design strategies and reliability assessments in engineering applications. Full article
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<p>Laminated specimen in experiments and numerical simulations.</p>
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<p>In-plane compressive load application diagram.</p>
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<p>The flow chart of experimental procedure.</p>
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<p>Experimental specimen and arrangement in this study.</p>
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<p>Convergence of FEA results of composite laminates specimen for different initial imperfection factors and mesh sizes. (<b>a</b>) Imperfection factor (IF); (<b>b</b>) mesh size (MS).</p>
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<p>Comparison between the post-buckling curves obtained from the experiment and FEM for the A—No delamination specimen.</p>
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<p>Comparison between the post-buckling curves obtained from the experiment and FEM for B—embedded delamination specimen with different delamination depth.</p>
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<p>Comparison between the post-buckling curves obtained from the experiment and FEM for B—embedded delamination specimens with different delamination length.</p>
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<p>Comparison between the post-buckling deformation obtained from the experiment and FEM for B—embedded delamination specimen. (<b>a</b>) Experimental specimen E43; (<b>b</b>) FEM model F43.</p>
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<p>The influence of the delamination depth on post-buckling curve of delaminated specimen obtained from FEA analysis.</p>
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<p>The influence of the delamination length on post-buckling curve of delaminated specimen obtained from FEA analysis.</p>
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19 pages, 1743 KiB  
Article
New Buckling Curve for a Compressed Member with Cold-Formed Channel Cross-Section
by Bálint Vaszilievits-Sömjén and Ferenc Papp
Buildings 2024, 14(10), 3258; https://doi.org/10.3390/buildings14103258 (registering DOI) - 15 Oct 2024
Viewed by 304
Abstract
The verification of a column made from a lipped cold-formed channel section, subjected to pure axial compression relative to the gross cross-section, often results in a combined verification of bending and compression due to the appearance of a shift of the centroid of [...] Read more.
The verification of a column made from a lipped cold-formed channel section, subjected to pure axial compression relative to the gross cross-section, often results in a combined verification of bending and compression due to the appearance of a shift of the centroid of its effective cross-section. Following Eurocode 3 rules, this requires the determination of two distinct effective cross-sections and various interaction factors. This paper, based on an analytic approach, offers a modification to the actual buckling curve, based on Ayrton–Perry formulation, to include the second-order effects raised by the eventual shift of the effective centroid due to local buckling of the compressed web plate. This eliminates the need to use an interaction formula. The modified buckling curve is verified based on a GMNIA analysis performed on a numerical parametric model, which was previously validated by laboratory tests. In addition, the results are compared with strength results provided by appropriate Eurocode 3 formulas and AISI Direct Strength Method for global-local interaction and with classic experimental results. Full article
(This article belongs to the Section Building Structures)
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<p>Concentrically compressed column subject to parametric study.</p>
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<p>“Pure” global, local and distortional imperfection modes assumed for the parametric study.</p>
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<p>Comparison of predicted resistances, Abaqus vs. use of new buckling curve, ±10% error ranges.</p>
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<p>Comparison of predicted resistances, AISI DSM [<a href="#B12-buildings-14-03258" class="html-bibr">12</a>] vs. use of new buckling curve, ±10% error ranges.</p>
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<p>Simple supported column with C200x60x22x1.50 mm profile.</p>
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<p>The deformed shape at the ultimate load level.</p>
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16 pages, 4576 KiB  
Article
A Method for the Coefficient Superposition Buckling Bearing Capacity of Thin-Walled Members
by Bing Xu, Lang Wang, Qin Liu, Rui Wang, Bing Kong and Bo Xu
Buildings 2024, 14(10), 3236; https://doi.org/10.3390/buildings14103236 - 12 Oct 2024
Viewed by 342
Abstract
Axial compression tests were conducted on short rhombic tubes of different cross-sectional shapes. The deformation modes of the rhombic short tubes were obtained. To induce a finite element model with deformation modes consistent with the actual working conditions, buckling modes are introduced into [...] Read more.
Axial compression tests were conducted on short rhombic tubes of different cross-sectional shapes. The deformation modes of the rhombic short tubes were obtained. To induce a finite element model with deformation modes consistent with the actual working conditions, buckling modes are introduced into the model as the initial imperfections of the structure. However, the buckling modes resulting from finite element buckling analyses often do not meet the needs of actual crushing modes. A coefficient superposition method of solution is proposed to derive modal characteristics consistent with the actual deformation modes by linear superposition of the buckling modes. Through the study of three aspects of theory, test, and simulation, and the comparison and verification of this method with the simulation results of related literature, the results show that the indexes derived from this method are closer to the actual circumstances and are more expandable, which provides a reference for the project. Full article
(This article belongs to the Section Building Structures)
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<p>Specimens for axial compression test and parameter definition.</p>
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<p>Test instrument (CMT5305).</p>
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<p>(<b>A</b>) Flat specimen for tension test. (<b>B</b>) Dimensions of the flat specimen. (<b>C</b>) True stress–strain curve.</p>
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<p>Deformation modes, buckling modes (<b>left</b>), and force–displacement curves of all the specimens (<b>right</b>) under axial compression.</p>
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<p>(<b>A</b>) Finite element model. (<b>B</b>) Direction definition for the model. (<b>C</b>) Mesh subdivision for the model.</p>
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<p>All 1st-order buckling mode results: (<b>A</b>) <math display="inline"><semantics> <mrow> <msub> <mrow> <mi>θ</mi> </mrow> <mrow> <mn>1</mn> </mrow> </msub> <mo>=</mo> <mn>30</mn> <mo>°</mo> </mrow> </semantics></math> (<b>B</b>) <math display="inline"><semantics> <mrow> <msub> <mrow> <mi>θ</mi> </mrow> <mrow> <mn>1</mn> </mrow> </msub> <mo>=</mo> <mn>45</mn> <mo>°</mo> </mrow> </semantics></math>. (<b>C</b>) <math display="inline"><semantics> <mrow> <msub> <mrow> <mi>θ</mi> </mrow> <mrow> <mn>1</mn> </mrow> </msub> <mo>=</mo> <mn>60</mn> <mo>°</mo> </mrow> </semantics></math>. (<b>D</b>) <math display="inline"><semantics> <mrow> <msub> <mrow> <mi>θ</mi> </mrow> <mrow> <mn>1</mn> </mrow> </msub> <mo>=</mo> <mn>90</mn> <mo>°</mo> </mrow> </semantics></math>. (<b>E</b>) Comparison of buckling modes of side plates under different boundaries.</p>
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<p>Smooth amplitude curve of simulated loading.</p>
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<p>The positive and negative definitions of buckling shape coefficients <math display="inline"><semantics> <mrow> <msub> <mrow> <mi mathvariant="bold-italic">a</mi> </mrow> <mrow> <mi mathvariant="bold-italic">i</mi> <mi mathvariant="bold-italic">j</mi> </mrow> </msub> </mrow> </semantics></math>, <math display="inline"><semantics> <mrow> <msub> <mrow> <mi mathvariant="bold-italic">b</mi> </mrow> <mrow> <mi mathvariant="bold-italic">j</mi> </mrow> </msub> </mrow> </semantics></math>.</p>
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<p>Deformation modes of the finite model (<b>left</b>), experimental model (<b>middle</b>), and displacement–force curve (<b>right</b>) under compression after the introduction of specific defects: (<b>A</b>) <math display="inline"><semantics> <mrow> <mstyle mathvariant="bold"> <msub> <mrow> <mi mathvariant="bold-italic">θ</mi> </mrow> <mrow> <mn>1</mn> </mrow> </msub> <mo>=</mo> <mn>30</mn> <mo>°</mo> </mstyle> </mrow> </semantics></math>. (<b>B</b>) <math display="inline"><semantics> <mrow> <mstyle mathvariant="bold"> <msub> <mrow> <mi mathvariant="bold-italic">θ</mi> </mrow> <mrow> <mn>1</mn> </mrow> </msub> <mo>=</mo> <mn>45</mn> <mo>°</mo> </mstyle> </mrow> </semantics></math>. (<b>C</b>) <math display="inline"><semantics> <mrow> <mstyle mathvariant="bold"> <msub> <mrow> <mi mathvariant="bold-italic">θ</mi> </mrow> <mrow> <mn>1</mn> </mrow> </msub> <mo>=</mo> <mn>60</mn> <mo>°</mo> </mstyle> </mrow> </semantics></math>. (<b>D</b>) <math display="inline"><semantics> <mrow> <mstyle mathvariant="bold"> <msub> <mrow> <mi mathvariant="bold-italic">θ</mi> </mrow> <mrow> <mn>1</mn> </mrow> </msub> <mo>=</mo> <mn>90</mn> <mo>°</mo> </mstyle> </mrow> </semantics></math>.</p>
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<p>The ratio of the initial peak crushing force between the simulated value and the experimental value under different numbers of sides.</p>
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<p>Comparison of two naming methods.</p>
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13 pages, 6880 KiB  
Article
The Evolution of Dilatant Shear Bands in High-Pressure Die Casting for Al-Si Alloys
by Jingzhou Lu, Ewan Lordan, Yijie Zhang, Zhongyun Fan, Wanlin Wang and Kun Dou
Materials 2024, 17(20), 5001; https://doi.org/10.3390/ma17205001 - 12 Oct 2024
Viewed by 401
Abstract
Bands of interdendritic porosity and positive macrosegregation are commonly observed in pressure die castings, with previous studies demonstrating their close relation to dilatant shear bands in granular materials. Despite recent technological developments, the micromechanism governing dilatancy in the high-pressure die casting (HPDC) process [...] Read more.
Bands of interdendritic porosity and positive macrosegregation are commonly observed in pressure die castings, with previous studies demonstrating their close relation to dilatant shear bands in granular materials. Despite recent technological developments, the micromechanism governing dilatancy in the high-pressure die casting (HPDC) process for alloys between liquid and solid temperature regions is still not fully understood. To investigate the influence of fluid flow and the size of externally solidified crystals (ESCs) on the evolution of dilatant shear bands in HPDC, various filling velocities were trialled to produce HPDC samples of Al8SiMnMg alloys. This study demonstrates that crystal fragmentation is accompanied by a decrease in dilatational concentration, producing an indistinct shear band. Once crystal fragmentation stagnates, the enhanced deformation rate associated with a further increase in filling velocity (from 2.2 ms−1 to 4.6 ms−1) localizes dilatancy into a highly concentrated shear band. The optimal piston velocity is 3.6 ms−1, under which the average ESC size reaches the minimum, and the average yield stress and overall product of strength and elongation reach the maximum values of 144.6 MPa and 3.664 GPa%, respectively. By adopting the concept of force chain buckling in granular media, the evolution of dilatant shear bands in equiaxed solidifying alloys can be adequately explained based on further verification with DEM-type modeling in OpenFOAM. Three mechanisms for ESC-enhanced dilation are presented, elucidating previous reports relating the presence of ESCs to the subsequent shear band characteristics. By applying the physics of granular materials to equiaxed solidifying alloys, unique opportunities are presented for process optimization and microstructural modeling in HPDC. Full article
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<p>Casting region with gating system and 8 tensile samples of HPDC tensile.</p>
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<p>Shot profile highlighting filling velocities used to produce HPDC tensile specimens.</p>
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<p>Infrared image and the temperature readings during various cycles of the HPDC process.</p>
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<p>Dilatant shear bands observed in HPDC samples produced with filling velocities of 2.2 ms<sup>−1</sup> (<b>i</b>), 3.6 ms<sup>−1</sup> (<b>ii</b>) and 4.2 ms<sup>−1</sup> (<b>iii</b>). Typical macrostructure of etched samples from the center of the gage section are shown (left) and corresponding EDX maps highlighting the eutectic fraction are shown (right).</p>
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<p>Optical micrographs taken from zone A, showing how dilatancy varies with filling velocities of (<b>i</b>) 2.2 ms<sup>−1</sup>, (<b>ii</b>) 3.6 ms<sup>−1</sup>, and (<b>iii</b>) 4.2 ms<sup>−1</sup>. Outlined in (<b>ii</b>) is a potential force chain that has persisted through deformation. The bulk filling direction was out of the page.</p>
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<p>Typical high-contrast secondary electron SEM micrographs used to obtain average ESC and in-cavity solidified grain size (<a href="#materials-17-05001-t001" class="html-table">Table 1</a>) for filling velocities of (<b>i</b>) 2.2 ms<sup>−1</sup>, (<b>ii</b>) 3.6 ms<sup>−1</sup> and (<b>iii</b>) 4.2 ms<sup>−1</sup>.</p>
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<p>The yield stress, elongation and ultimate tensile stress of various tensile samples.</p>
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<p>Graphical illustration highlighting the three mechanisms governing ESC enhanced dilation within the shear band: (i) “Stacking faults” introduced by the presence of ESCs along the force chain; (ii) ESCs located on the outermost regions of the band effectively acting as pivots; (iii) ESCs propelled by highly turbulent flow conditions, potentially dislodging crystals from the force chain. σ denotes the major principle stress axis.</p>
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<p>(<b>i</b>) Calculation domain and mesh for the model. (<b>ii</b>) The fluid velocity distribution of melt. (<b>iii</b>) The ESC motion and velocity distribution in the tensile sample during filling.</p>
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<p>The ESCs’ aggregation tendency in the tensile sample during filling.</p>
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18 pages, 7020 KiB  
Article
Axial Impact Response of Carbon Fiber-Reinforced Polymer Structures in High-Speed Trains Based on Filament Winding Process
by Aiqin Tian, Kang Sun, Quanwei Che, Beichen Jiang, Xiangang Song, Lirong Guo, Dongdong Chen and Shoune Xiao
Materials 2024, 17(20), 4970; https://doi.org/10.3390/ma17204970 - 11 Oct 2024
Viewed by 379
Abstract
The continuous increase in the operating speed of rail vehicles demands higher requirements for passive safety protection and lightweight design. This paper focuses on an energy-absorbing component (circular tubes) at the end of a train. Thin-walled carbon fiber-reinforced polymer (CFRP) tubes were prepared [...] Read more.
The continuous increase in the operating speed of rail vehicles demands higher requirements for passive safety protection and lightweight design. This paper focuses on an energy-absorbing component (circular tubes) at the end of a train. Thin-walled carbon fiber-reinforced polymer (CFRP) tubes were prepared using the filament winding process. Through a combination of sled impact tests and finite element simulations, the effects of a chamfered trigger (Tube I) and embedded trigger (Tube II) on the impact response and crashworthiness of the structure were investigated. The results showed that both triggering methods led to the progressive end failure of the tubes. Tube I exhibited a mean crush force (MCF) of 891.89 kN and specific energy absorption (SEA) of 38.69 kJ/kg. In comparison, the MCF and SEA of Tube II decreased by 21.2% and 21.9%, respectively. The reason for this reduction is that the presence of the embedded trigger in Tube II restricts the expansion of the inner plies (plies 4 to 6), thereby affecting the overall energy absorption mechanism. Based on the validated finite element model, a modeling strategy study was conducted, including the failure parameters (DFAILT/DFAILC), the friction coefficient, and the interfacial strength. It was found that the prediction results are significantly influenced by modeling methods. Specifically, as the interfacial strength decreases, the tube wall is more prone to circumferential cracking or overall buckling under axial impact. Full article
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<p>Preparation of composite material tubes.</p>
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<p>Dynamic impact test.</p>
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<p>Numerical model for axial crushing of CFRP tubes.</p>
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<p>Schematic of delamination failure mode.</p>
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<p>Axial crushing force–displacement curves from experiment and simulation: (<b>a</b>) Tube I; (<b>b</b>) Tube II.</p>
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<p>Crushing process of Tube I from experiment and simulation.</p>
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<p>Crushing process of Tube II from experiment and simulation.</p>
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<p>Cross-sectional view of tube wall during crushing: (<b>a</b>) Tube I; (<b>b</b>) Tube II.</p>
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<p>Damage distribution of composite material tubes.</p>
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<p>Comparison of experimental and simulated crushing responses of Tube I under different fiber failure parameters: (<b>a</b>) impact force–displacement curves; (<b>b</b>) energy absorption–displacement curves.</p>
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<p>Effects of failure parameters (DFAILT/DFAILC), friction coefficient, and interfacial strength on MCF and SEA.</p>
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<p>Crushing process of Tube I predicted using different failure parameters.</p>
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<p>Comparison of experimental and simulated crushing responses of Tube I under different friction coefficients: (<b>a</b>) impact force–displacement curves; (<b>b</b>) energy absorption–displacement curves.</p>
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<p>Simulated crushing process of Tube I under different friction coefficients.</p>
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<p>Comparison of experimental and simulated crushing responses of Tube I under different inter-ply failure stresses: (<b>a</b>) crushing force–displacement curves; (<b>b</b>) energy absorption–displacement curves.</p>
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<p>Simulated crushing process of Tube I under different inter-ply failure stresses.</p>
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16 pages, 3480 KiB  
Article
Evaluating the Seismic Resilience of Above-Ground Liquid Storage Tanks
by Emanuele Brunesi and Roberto Nascimbene
Buildings 2024, 14(10), 3212; https://doi.org/10.3390/buildings14103212 - 9 Oct 2024
Viewed by 407
Abstract
Historical seismic events have repeatedly highlighted the susceptibility of above-ground liquid storage steel tanks, underscoring the critical need for their proper design to minimize potential damage due to seismic forces. A significant failure mechanism in these structures, which play essential roles in the [...] Read more.
Historical seismic events have repeatedly highlighted the susceptibility of above-ground liquid storage steel tanks, underscoring the critical need for their proper design to minimize potential damage due to seismic forces. A significant failure mechanism in these structures, which play essential roles in the extraction and distribution of various raw or refined materials—many of which are flammable or environmentally hazardous—is the dynamic buckling of the tank walls. This study introduces a numerical framework designed to assess the earthquake-induced hydrodynamic pressures exerted on the walls of cylindrical steel tanks. These pressures result from the inertial forces generated during seismic activity. The computational framework incorporates material and geometric nonlinearities and models the tanks using four-node shell elements with two-point integration, specifically Belytschko shell elements. The Arbitrary Lagrangian–Eulerian (ALE) method is employed to accommodate substantial structural and fluid deformations, enabling a full simulation of fluid–structure interaction through highly nonlinear algorithms. Experimental test data are utilized to validate the proposed modeling approach, particularly in replicating sloshing phenomena and identifying stress concentrations that may lead to wall buckling. The study further presents results from a parametric analysis that varies the height-to-radius and radius-to-thickness ratios of a typical anchored flat-bottomed tank, examining the seismic performance of this common storage system. These results provide insights into the relationship between tank properties and mechanical behavior under dynamic loading conditions. Full article
(This article belongs to the Section Building Structures)
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<p>(<b>a</b>–<b>c</b>) illustrate examples of elephant’s foot buckling mechanisms, while (<b>d</b>–<b>f</b>) show instances of diamond-shaped buckling. These failure modes were commonly observed in the aftermath of the May 2012 Emilia seismic sequence.</p>
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<p>Anchoring system-related failure modes were observed in both flat-base and leg-supported storage tanks during the May 2012 Emilia seismic sequence. (<b>a</b>,<b>b</b>) demonstrate issues in flat-base tanks, including the fracture of anchor bolts and significant spalling of the concrete foundation at the anchorage points; (<b>c</b>,<b>d</b>) depict leg-supported tanks that suffered from a loss of verticality, primarily due to shear-buckling in the stocky, tapered support legs.</p>
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<p>The high-definition finite element (FE) model was meticulously developed to closely simulate the outcomes of shake-table tests. This advanced modeling approach was essential for accurately replicating the dynamic responses and structural behaviors observed during controlled seismic testing, providing vital insights into the performance of the structures under simulated earthquake conditions.</p>
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<p>The tank specimen evaluated by Manos and Clough [<a href="#B55-buildings-14-03212" class="html-bibr">55</a>] is a scaled model, specifically designed to facilitate detailed study under laboratory conditions: summary of the main geometrical properties.</p>
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<p>The comparison of pressure time histories from numerical simulations and experimental tests conducted by Manos and Clough [<a href="#B55-buildings-14-03212" class="html-bibr">55</a>] reveals a meticulous alignment of results, affirming the robustness of the numerical models used.</p>
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<p>Finite element model showcases the deformed shapes of the tank at different time instants, visually demonstrating how the structure distorts in response to seismic forces (as tested by Haroun [<a href="#B56-buildings-14-03212" class="html-bibr">56</a>]).</p>
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<p>Elastic acceleration displacement response spectra from recorded time histories at MRN station (NS and EW) and in accordance with the Italian building code [<a href="#B62-buildings-14-03212" class="html-bibr">62</a>].</p>
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<p>Total wall pressure distribution along the height for (<b>a</b>) <math display="inline"><semantics> <mrow> <mi>H</mi> <mo>/</mo> <mi>R</mi> </mrow> </semantics></math> = 1.5 and (<b>b</b>) <math display="inline"><semantics> <mrow> <mi>H</mi> <mo>/</mo> <mi>R</mi> </mrow> </semantics></math> = 3.0.</p>
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<p>Elastic displacement response spectra from ground motion records selected by Maley et al. [<a href="#B63-buildings-14-03212" class="html-bibr">63</a>].</p>
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<p>Wall pressure peak profiles (individual records, mean, and mean plus one standard deviation) for (<b>a</b>) <math display="inline"><semantics> <mrow> <mi>H</mi> <mo>/</mo> <mi>R</mi> </mrow> </semantics></math> = 0.75 and <math display="inline"><semantics> <mrow> <mi>R</mi> <mo>/</mo> <mi>t</mi> </mrow> </semantics></math> = 2000, (<b>b</b>) <math display="inline"><semantics> <mrow> <mi>H</mi> <mo>/</mo> <mi>R</mi> </mrow> </semantics></math> = 1.5 and <math display="inline"><semantics> <mrow> <mi>R</mi> <mo>/</mo> <mi>t</mi> </mrow> </semantics></math> = 400, and (<b>c</b>) <math display="inline"><semantics> <mrow> <mi>H</mi> <mo>/</mo> <mi>R</mi> </mrow> </semantics></math> = 4 and <math display="inline"><semantics> <mrow> <mi>R</mi> <mo>/</mo> <mi>t</mi> </mrow> </semantics></math> = 400.</p>
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<p>Comparison between wall pressure peak profiles (mean and mean plus standard deviation).</p>
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23 pages, 7073 KiB  
Article
Risk Assessment of Overturning of Freestanding Non-Structural Building Contents in Buckling-Restrained Braced Frames
by Atsushi Suzuki, Susumu Ohno and Yoshihiro Kimura
Buildings 2024, 14(10), 3195; https://doi.org/10.3390/buildings14103195 - 8 Oct 2024
Viewed by 493
Abstract
The increasing demand in structural engineering now extends beyond collapse prevention to encompass business continuity planning (BCP). In response, energy dissipation devices have garnered significant attention for building response control. Among these, buckling-restrained braces (BRBs) are particularly favored due to their stable hysteretic [...] Read more.
The increasing demand in structural engineering now extends beyond collapse prevention to encompass business continuity planning (BCP). In response, energy dissipation devices have garnered significant attention for building response control. Among these, buckling-restrained braces (BRBs) are particularly favored due to their stable hysteretic behavior and well-established design provisions. However, BCP also necessitates the prevention of furniture overturning—an area that remains quantitatively underexplored in the context of buckling-restrained braced frames (BRBFs). Addressing this gap, this research designs BRBFs using various design criteria and performs incremental dynamic analysis (IDA) with artificially generated seismic waves. The results are compared with previously developed fragility curves for furniture overturning under different BRB design conditions. The findings demonstrate that the fragility of furniture overturning can be mitigated by a natural frequency shift, which alters the threshold of critical peak floor acceleration. These results, combined with hazard curves obtained from various locations across Japan, quantify the mean annual frequency of furniture overturning. The study reveals that increased floor acceleration in stiffer BRBFs can lead to a 3.8-fold higher risk of furniture overturning compared to frames without BRBs. This heightened risk also arises from the greater hazards at shorter natural periods due to stricter response reduction demands. The probabilistic risk analysis, which integrates fragility and hazard assessments, provides deeper insights into the evaluation of BCP. Full article
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<p>Configuration of Buckling-Restrained Braced Frame (BRBF): (<b>a</b>) diagonal configuration; (<b>b</b>) V configuration.</p>
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<p>Concept of this research: (<b>a</b>) MRF; (<b>b</b>) BRBF; (<b>c</b>) furniture overturning after the 2022 Fukushima-Oki earthquake.</p>
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<p>Drawing and member schedule of model structure.</p>
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<p>Idealized response spectra: (<b>a</b>) displacement; (<b>b</b>) velocity; (<b>c</b>) acceleration.</p>
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<p>Calibration of BRB specification: (<b>a</b>) displacement response spectrum; (<b>b</b>) performance curve.</p>
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<p>FEA model of model structure.</p>
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<p>Fragility curve of furniture overturning.</p>
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<p>Earthquake scenarios concerned.</p>
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<p>Acceleration response spectra (<span class="html-italic">h</span> = 0.02).</p>
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<p>Maximum inter-story drift obtained from IDA: (<b>a</b>) MRF; (<b>b</b>) <span class="html-italic">θ<sub>t</sub> =</span> 1/120 rad; (<b>c</b>) <span class="html-italic">θ<sub>t</sub> =</span> 1/150 rad; (<b>d</b>) <span class="html-italic">θ<sub>t</sub> =</span> 1/200 rad (<span class="html-italic">γ</span><sub>I</sub><span class="html-italic">a<sub>g</sub></span> = 3.0 m/s<sup>2</sup>); (<b>e</b>) <span class="html-italic">θ<sub>t</sub> =</span> 1/200 rad (<span class="html-italic">γ</span><sub>I</sub><span class="html-italic">a<sub>g</sub></span> = 3.6 m/s<sup>2</sup>); (<b>f</b>) <span class="html-italic">θ<sub>t</sub> </span>= 1/200 rad (<span class="html-italic">γ</span><sub>I</sub><span class="html-italic">a<sub>g</sub></span> = 4.2 m/s<sup>2</sup>).</p>
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<p>Peak floor acceleration obtained from IDA: (<b>a</b>) MRF; (<b>b</b>) <span class="html-italic">θ<sub>t</sub> =</span> 1/120 rad; (<b>c</b>) <span class="html-italic">θ<sub>t</sub> =</span> 1/150 rad; (<b>d</b>) <span class="html-italic">θ<sub>t</sub> =</span> 1/200 rad (<span class="html-italic">γ</span><sub>I</sub><span class="html-italic">a<sub>g</sub></span> = 3.0 m/s<sup>2</sup>); (<b>e</b>) <span class="html-italic">θ<sub>t</sub> =</span> 1/200 rad (<span class="html-italic">γ</span><sub>I</sub><span class="html-italic">a<sub>g</sub></span> = 3.6 m/s<sup>2</sup>); (<b>f</b>) <span class="html-italic">θ<sub>t</sub> </span>= 1/200 rad (<span class="html-italic">γ</span><sub>I</sub><span class="html-italic">a<sub>g</sub></span> = 4.2 m/s<sup>2</sup>).</p>
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<p>Threshold acceleration causing 50% of furniture overturning (low): (<b>a</b>) MRF; (<b>b</b>) <span class="html-italic">θ<sub>t</sub> =</span> 1/120 rad; (<b>c</b>) <span class="html-italic">θ<sub>t</sub> =</span> 1/150 rad; (<b>d</b>) <span class="html-italic">θ<sub>t</sub> =</span> 1/200 rad (<span class="html-italic">γ</span><sub>I</sub><span class="html-italic">a<sub>g</sub></span> = 3.0 m/s<sup>2</sup>); (<b>e</b>) <span class="html-italic">θ<sub>t</sub> =</span> 1/200 rad (<span class="html-italic">γ</span><sub>I</sub><span class="html-italic">a<sub>g</sub></span> = 3.6 m/s<sup>2</sup>); (<b>f</b>) <span class="html-italic">θ<sub>t</sub> =</span> 1/200 rad (<span class="html-italic">γ</span><sub>I</sub><span class="html-italic">a<sub>g</sub></span> = 4.2 m/s<sup>2</sup>).</p>
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<p>Calculation flow of fragility curve of furniture overturning.</p>
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<p>Fragility curve of furniture overturning: (<b>a</b>) low; (<b>b</b>) medium; (<b>c</b>) tall.</p>
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<p>Computation concept of mean annual frequency: (<b>a</b>) fragility curve; (<b>b</b>) derivative of hazard curve; (<b>c</b>) mean annual frequency.</p>
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<p>Hazard curves for locations across Japan: (<b>a</b>) Hokkaido; (<b>b</b>) Miyagi; (<b>c</b>) Tokyo; (<b>d</b>) Ishikawa; (<b>e</b>) Aichi; (<b>f</b>) Hyogo; (<b>g</b>) Hiroshima; (<b>h</b>) Kumamoto.</p>
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<p>Mean annual frequency of furniture overturning: (<b>a</b>) low; (<b>b</b>) medium; (<b>c</b>) tall.</p>
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13 pages, 2572 KiB  
Article
Investigation of the Bending Process and Theory in Free-Boundary Pneumatic Film-Forming for Curved Image Sensors
by Weihan Zheng, Chunlai Li, Jiangcheng Hu and Liang Guo
Sensors 2024, 24(19), 6428; https://doi.org/10.3390/s24196428 - 4 Oct 2024
Viewed by 440
Abstract
To explore the bending process and theory of the free-boundary aerodynamic film forming method for curved detectors, this study integrates practical forming structures with theoretical analysis and establishes a simulation model to investigate stress, strain, and morphological changes during bending. The analysis indicates [...] Read more.
To explore the bending process and theory of the free-boundary aerodynamic film forming method for curved detectors, this study integrates practical forming structures with theoretical analysis and establishes a simulation model to investigate stress, strain, and morphological changes during bending. The analysis indicates that the shift from “projection” to “wrapping” in forming theory is due to the release of boundary degrees of freedom. The forming process can be summarized as the mold’s arc characteristics, originating from the chip’s corners, gradually replacing the chip’s rectangular characteristics along the diagonal, resulting in corresponding stress and strain changes. The “wrapping” bending theory of this method has significant advantages over traditional methods and represents a crucial direction for achieving higher curvature in the future. However, this study found that the use of film pressure can only inhibit out-of-plane deformation to a certain extent, and the buckling phenomenon will still occur when the thinner chip is bent. It prevents the use of thinner chips in the thinning–bending method, so avoiding out-of-plane deformation during the molding process is the direction that needs to be broken in the future. Full article
(This article belongs to the Section Sensing and Imaging)
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<p>Principle comparison of imaging methods: (<b>a</b>) distortion principle of traditional image sensor due to Petzval field curve; (<b>b</b>) imaging principle of curved image sensor inspired by human eye.</p>
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<p>Forming principle; the left side is the traditional fixed-boundary mold forming, while the right side is the free-boundary pneumatic film forming.</p>
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<p>When using free-boundary mold molding, the molding method has been changed from “projection” to “wrap”.</p>
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<p>If the difference between the displacement in the x and y direction and the displacement in diagonal direction is too large, there will be a tendency of out-of-plane deformation.</p>
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<p>Points x, y and d in the figure represent the intersection of the dividing line and the lines in the x, y and d directions. Make the x, y, and d points on the fitting boundary located on the same section diagram. The points representing the diagonal direction on the fitting boundary gradually approach and exceed the x and y points during the forming process. It indicates that the fitting speed in the d direction is faster than that in the x and y directions.</p>
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<p>Structure diagram of free-boundary pneumatic film forming method.</p>
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<p>Edge fitting direction.</p>
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<p>At the beginning of bending, the Mises stress was mainly affected by the length of the chip, and the Mises stress on the <span class="html-italic">y</span> axis increased to distinguish the low stress in the middle of the chip. (<b>a</b>) Before the division of the middle low-stress region; (<b>b</b>) After the segmentation of the low stress region in the middle.</p>
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<p>Diagram of stress distribution when edge fitting is completed (S11 represents radial stress, and S22 represents circumferential stress): (<b>a</b>) the shaded area indicates where bonding is complete, forming an arcuate boundary line near the edges. Stress concentration occurs in the central part of the unbonded arc; (<b>b</b>) circumferential stress distribution after edge fitting.</p>
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<p>Buckling from the diagonal caused by the chip being too thin.</p>
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<p>Repeated changes of the chip edge midpoint displacement and subsequent laminating progress reduce the edge fitting speed. (<b>a</b>) Displacement changes during normal forming process; (<b>b</b>) The change of displacement at the midpoint of the edge before buckling occurs.</p>
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<p>Three forming directions in the forming stage.</p>
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<p>Affected by the direction of the diagonal fit, the shape of the fit boundary changes: (<b>a</b>) after the end of edge fitting for a period of time, the arc boundary line from the four corners gradually replaces the rectangular boundary line; (<b>b</b>) boundary shape near the end of molding.</p>
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<p>Mises stress distribution during molding process: (<b>a</b>) Mises stress distribution at the beginning of the molding process; (<b>b</b>) Mises stress distribution during molding; (<b>c</b>) Mises stress distribution change during molding process.</p>
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<p>On the boundary point d passes point x and does not reach point y: (<b>a</b>) The circular boundary line from the four corners has just swallowed up the short boundary line diagram; (<b>b</b>) at this time, the compressive stress in this region reaches its peak, and the maximum stress and range in this region are smaller than that in the concentrated region near the edge; (<b>c</b>) at this time, the area indicated in Figure (<b>c</b>) has been fitted and there is no longer any stress concentration. Taking the region c in the Figure as an example, it is shown that the stress will produce a concentrated phenomenon of regional stress change process.</p>
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<p>In the center of the chip appears a droplet area symmetrical about the <span class="html-italic">y</span>-axis.</p>
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<p>Radial stress distribution at the end of molding.</p>
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<p>Strain distribution in the chip after molding: (<b>a</b>) radial strain results; (<b>b</b>) tangential strain results.</p>
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17 pages, 16700 KiB  
Article
Experimental Study and Design Method of Cold-Formed Thin-Walled Steel Unequal-Leg Angles under Axial Compression
by Yanli Guo, Zeyu Nie, Xingyou Yao, Yilin Liu, Chong Chen and Kaihua Zeng
Buildings 2024, 14(10), 3132; https://doi.org/10.3390/buildings14103132 - 30 Sep 2024
Viewed by 406
Abstract
An experimental study of cold-formed thin-walled steel unequal-leg angles (CFTWS-ULAs) under axially oriented pressure is presented in this paper. Firstly, the initial imperfections and material properties of the angle specimens were measured in detail. The angle specimens were tested under fixed-ended conditions. The [...] Read more.
An experimental study of cold-formed thin-walled steel unequal-leg angles (CFTWS-ULAs) under axially oriented pressure is presented in this paper. Firstly, the initial imperfections and material properties of the angle specimens were measured in detail. The angle specimens were tested under fixed-ended conditions. The results of the experiments showed that the angle specimens with small slenderness ratios were susceptible to local buckling, the angle specimens whose legs had high slenderness ratios and low width–thickness ratios were found to easily suffer from the occurrence of flexural buckling, and the angle specimens whose legs had high width–thickness ratios were found to easily suffer from the occurrence of interactive buckling between local buckling and flexural buckling. The finite element analysis of the ULAs was conducted using ABAQUS6.14 finite element software by creating a model. The buckling modes and ultimate bearing capacities of the test specimens were compared, and the finite element analysis verified that the established model built using the finite element is credible and subsequent parametric analysis was performed. The slenderness ratio had the most significant impact on the ultimate bearing capacities of the unequal-leg angles, as indicated by the analysis results. When the width–thickness ratio and the width ratio of the legs fell within a specific range, the ultimate bearing capacities of the unequal-leg angles increased as the width–thickness ratio and the width ratio of the legs increased. Finally, the comparison results showed that the design strengths predicted by the specifications were very conservative, because the local buckling and torsional buckling were calculated at the same time. Therefore, a recommendation was proposed that the calculation of the load-carrying capacity of an unequal-leg angle should ignore torsional buckling. Full article
(This article belongs to the Special Issue Structural Performance of Building Steel)
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<p>Cross-section of unequal-leg angle.</p>
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<p>Specimen numbering rules.</p>
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<p>Stress–strain curves.</p>
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<p>The measurement position of longitudinal initial geometric imperfections.</p>
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<p>Initial geometric imperfections of CFTWS-ULA specimens.</p>
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<p>Device set-up.</p>
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<p>Displacement meter layout.</p>
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<p>Buckling types of the 400 mm CFTWS-ULA specimens.</p>
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<p>Buckling types of the 900 mm CFTWS-ULA specimens.</p>
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<p>Buckling types of the 1500 mm CFTWS-ULA specimens.</p>
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<p>Buckling types of the 2100 mm CFTWS-ULA specimens.</p>
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<p>Test load–axial displacement curves of CFTWS-ULAs.</p>
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<p>Test load–axial displacement curves of CFTWS-ULAs.</p>
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<p>Finite element model of CFTWS-ULA specimen.</p>
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<p>Comparison of buckling types between experiment and finite element analysis.</p>
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<p>Comparison of load–displacement curves between test and finite element analysis.</p>
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<p>Curves between slenderness ratios and ultimate bearing capacities.</p>
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<p>Curves between width ratio of legs and ultimate bearing capacities.</p>
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<p>Curves of width-to-thickness ratios and ultimate bearing capacities.</p>
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23 pages, 3108 KiB  
Article
Thermal Buckling and Postbuckling Analysis of Cracked FG-GPL RC Plates Using a Phase-Field Crack Model
by Jin-Rae Cho
Appl. Sci. 2024, 14(19), 8794; https://doi.org/10.3390/app14198794 - 29 Sep 2024
Viewed by 475
Abstract
A phase-field crack model is developed for numerical analysis of thermal buckling and postbuckling behavior of a functionally graded (FG) graphene platelet-reinforced composite (FG-GPLRC) plate with a central crack. The inclined central crack is represented according to a stable, effective phase-field formulation (PFF) [...] Read more.
A phase-field crack model is developed for numerical analysis of thermal buckling and postbuckling behavior of a functionally graded (FG) graphene platelet-reinforced composite (FG-GPLRC) plate with a central crack. The inclined central crack is represented according to a stable, effective phase-field formulation (PFF) by introducing a virtual crack rotation. The problem is formulated using first-order shear deformation theory (SDT) incorporated with von Kármán geometric nonlinearity. And it is approximated by combining regular Laplace interpolation functions and crack-tip singular functions in the framework of the 2D extended natural element method (XNEM). Troublesome shear locking is suppressed by applying the concept of the MITC (mixed-interpolated tensorial components)3+ shell element to the present numerical method. The results demonstrate the effectiveness of this method in accurately predicting the critical buckling temperature rise (CBTR) and the thermal postbuckling path. In addition, the parametric results reveal that the CBTR and postbuckling path of the FG-GPLRC plate are distinct from those of the FG carbon nanotube-reinforced composite (FG-CNTRC) plate and remarkably affected by the parameters associated with the crack and graphene platelet (GPL). Full article
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<p>Graphene platelet-reinforced rectangular plate: (<b>a</b>) Geometry and dimensions; (<b>b</b>) Primitive GPL distribution patterns.</p>
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<p>The phase field <math display="inline"><semantics> <mrow> <mi>ψ</mi> <mfenced> <mstyle mathvariant="bold-italic" mathsize="normal"> <mi>x</mi> </mstyle> </mfenced> </mrow> </semantics></math> within the cracked FG-GPLRC plate.</p>
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<p>Uniform NEM grid and L/I functions.</p>
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<p>Virtual rotation of inclined crack in the PFF.</p>
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<p>A flowchart for thermal buckling and postbucking analysis of the cracked FG-PLRC plate.</p>
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<p>Comparison of thermal postbuckling paths of a PIMM simply-supported isotropic plate under uniform temperature rise [<a href="#B19-applsci-14-08794" class="html-bibr">19</a>,<a href="#B54-applsci-14-08794" class="html-bibr">54</a>].</p>
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<p>CBTR <math display="inline"><semantics> <mrow> <mo>Δ</mo> <msub> <mi>T</mi> <mrow> <mi>c</mi> <mi>r</mi> </mrow> </msub> <mfenced> <mi>K</mi> </mfenced> </mrow> </semantics></math> of the intact FG-GPLRC plate (SSSS): (<b>a</b>) The relative thickness ratio <math display="inline"><semantics> <mrow> <mi>a</mi> <mo>/</mo> <mi>h</mi> </mrow> </semantics></math> (<math display="inline"><semantics> <mrow> <msubsup> <mi>g</mi> <mrow> <mi>G</mi> <mi>P</mi> <mi>L</mi> </mrow> <mo>*</mo> </msubsup> <mo>=</mo> <mn>0.3</mn> <mo>%</mo> </mrow> </semantics></math>); (<b>b</b>) The total GPL mass.<math display="inline"><semantics> <mrow> <msubsup> <mi>g</mi> <mrow> <mi>G</mi> <mi>P</mi> <mi>L</mi> </mrow> <mo>*</mo> </msubsup> </mrow> </semantics></math> (<math display="inline"><semantics> <mrow> <mi>a</mi> <mo>/</mo> <mi>h</mi> <mo>=</mo> <mn>25</mn> </mrow> </semantics></math>, PIMM).</p>
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<p>CBTR <math display="inline"><semantics> <mrow> <mo>Δ</mo> <msub> <mi>T</mi> <mrow> <mi>c</mi> <mi>r</mi> </mrow> </msub> <mfenced> <mi>K</mi> </mfenced> </mrow> </semantics></math> of the cracked FG-GPLRC plate (FG-U, <math display="inline"><semantics> <mrow> <msubsup> <mi>g</mi> <mrow> <mi>G</mi> <mi>P</mi> <mi>L</mi> </mrow> <mo>*</mo> </msubsup> <mo>=</mo> <mn>0.3</mn> <mo>%</mo> </mrow> </semantics></math>): (<b>a</b>) The crack length <math display="inline"><semantics> <mrow> <mi>c</mi> <mo>/</mo> <mi>a</mi> </mrow> </semantics></math> (<math display="inline"><semantics> <mrow> <mi>α</mi> <mo>=</mo> <msup> <mn>0</mn> <mo>°</mo> </msup> </mrow> </semantics></math>); (<b>b</b>) The crack angle <math display="inline"><semantics> <mi>α</mi> </semantics></math> (<math display="inline"><semantics> <mrow> <mi>c</mi> <mo>/</mo> <mi>a</mi> <mo>=</mo> <mn>0.3</mn> </mrow> </semantics></math>, FIMM).</p>
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<p>Variation in thermal postbuckling paths (FG-U, <math display="inline"><semantics> <mrow> <msubsup> <mi>g</mi> <mrow> <mi>G</mi> <mi>P</mi> <mi>L</mi> </mrow> <mo>*</mo> </msubsup> <mo>=</mo> <mn>0.3</mn> <mo>%</mo> </mrow> </semantics></math>, <math display="inline"><semantics> <mrow> <mi>a</mi> <mo>/</mo> <mi>h</mi> <mo>=</mo> <mn>25</mn> </mrow> </semantics></math>, CCCC): (<b>a</b>) The crack relative length <math display="inline"><semantics> <mrow> <mi>c</mi> <mo>/</mo> <mi>a</mi> </mrow> </semantics></math> (<math display="inline"><semantics> <mrow> <mi>α</mi> <mo>=</mo> <msup> <mn>0</mn> <mo>°</mo> </msup> </mrow> </semantics></math>); (<b>b</b>) The crack angle (<math display="inline"><semantics> <mrow> <mi>c</mi> <mo>/</mo> <mi>a</mi> <mo>=</mo> <mn>0.3</mn> </mrow> </semantics></math>).</p>
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<p>Variation in thermal postbuckling paths (<math display="inline"><semantics> <mrow> <mi>c</mi> <mo>/</mo> <mi>a</mi> <mo>=</mo> <mn>0.3</mn> <mo>,</mo> <mo> </mo> <mi>α</mi> <mo>=</mo> <msup> <mrow> <mn>30</mn> </mrow> <mo>°</mo> </msup> </mrow> </semantics></math>, CCCC): (<b>a</b>) The GPL mass fraction <math display="inline"><semantics> <mrow> <msubsup> <mi>g</mi> <mrow> <mi>G</mi> <mi>P</mi> <mi>L</mi> </mrow> <mo>*</mo> </msubsup> </mrow> </semantics></math> (FG-O); (<b>b</b>) The GPL distribution pattern (<math display="inline"><semantics> <mrow> <msubsup> <mi>g</mi> <mrow> <mi>G</mi> <mi>P</mi> <mi>L</mi> </mrow> <mo>*</mo> </msubsup> <mo>=</mo> <mn>0.2</mn> <mo>%</mo> </mrow> </semantics></math>).</p>
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<p>Variation in thermal postbuckling paths (FG-U, <math display="inline"><semantics> <mrow> <msubsup> <mi>g</mi> <mrow> <mi>G</mi> <mi>P</mi> <mi>L</mi> </mrow> <mo>*</mo> </msubsup> <mo>=</mo> <mn>0.3</mn> <mo>,</mo> <mo> </mo> <mi>c</mi> <mo>/</mo> <mi>a</mi> <mo>=</mo> <mn>0.3</mn> <mo>,</mo> <mo> </mo> <mi>α</mi> <mo>=</mo> <msup> <mrow> <mn>45</mn> </mrow> <mo>°</mo> </msup> </mrow> </semantics></math>): (<b>a</b>) The aspect ratio <math display="inline"><semantics> <mrow> <mi>b</mi> <mo>/</mo> <mi>a</mi> </mrow> </semantics></math> (CCCC); (<b>b</b>) The boundary condition.</p>
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<p>CNTRC (FG-U, <math display="inline"><semantics> <mrow> <mi>c</mi> <mo>/</mo> <mi>a</mi> <mo>=</mo> <mn>0.3</mn> </mrow> </semantics></math>, CCCC): (<b>a</b>) Critical buckling temperature rise <math display="inline"><semantics> <mrow> <mo>Δ</mo> <msub> <mi>T</mi> <mrow> <mi>c</mi> <mi>r</mi> </mrow> </msub> <mfenced> <mi>K</mi> </mfenced> </mrow> </semantics></math> (<math display="inline"><semantics> <mrow> <msubsup> <mi>V</mi> <mrow> <mi>c</mi> <mi>n</mi> <mi>t</mi> </mrow> <mo>*</mo> </msubsup> <mo>=</mo> <mn>0.12</mn> </mrow> </semantics></math>); (<b>b</b>) Comparison of postbuckling paths with GLRC (<math display="inline"><semantics> <mrow> <mi>α</mi> <mo>=</mo> <msup> <mn>0</mn> <mo>°</mo> </msup> <mo>,</mo> <msub> <mi>α</mi> <mrow> <mi>c</mi> <mi>n</mi> <mi>t</mi> </mrow> </msub> <mo>=</mo> <msup> <mrow> <mn>90</mn> </mrow> <mo>°</mo> </msup> </mrow> </semantics></math>).</p>
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<p>Six tying points within a three-node master element <math display="inline"><semantics> <mover accent="true"> <mi>ϖ</mi> <mo stretchy="false">^</mo> </mover> </semantics></math>.</p>
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32 pages, 775 KiB  
Review
A Comprehensive Synthesis on Analytical Algorithms for Assessing Elastic Buckling Loads of Thin-Walled Isotropic and Laminated Cylindrical Shells
by Maria Tănase
Processes 2024, 12(10), 2120; https://doi.org/10.3390/pr12102120 - 29 Sep 2024
Viewed by 297
Abstract
A comprehensive review is presented on the main analytical methods used in the specialized literature to evaluate the buckling loads of thin-walled cylindrical shells (TWCS) subjected to different mechanical loads or load combinations. The analytical formulations are first presented for unstiffened TWCS, followed [...] Read more.
A comprehensive review is presented on the main analytical methods used in the specialized literature to evaluate the buckling loads of thin-walled cylindrical shells (TWCS) subjected to different mechanical loads or load combinations. The analytical formulations are first presented for unstiffened TWCS, followed by stiffened TWCS in different configurations (stiffeners in the axial direction, circumferential direction or both axial and circumferential directions, placed on the external or internal surface of the shell). This research can serve as a helpful resource for researchers investigating this field, allowing the analytical methods to be used as a reference basis for numerical and experimental results regarding the behavior of structures in the category of TWCS. Full article
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Figure 1
<p>Evaluation scheme for the buckling behavior of thin cylindrical shells according to their constructive characteristics.</p>
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<p>Methodology of the research.</p>
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<p>The geometry of a laminate with N layers.</p>
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<p>The coordinate system of the laminate: (<b>a</b>) fibers parallel to the x-axis of the global coordinate system, (<b>b</b>) fibers oriented in a direction forming a certain angle with the x-axis of the global coordinate system.</p>
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25 pages, 16751 KiB  
Article
Optimization of Shear Resistance in Horizontal Joints of Prefabricated Shear Walls through Post-Cast Epoxy Resin Concrete Applications
by Peiqi Chen, Shilong Zhao, Pengzhan Xu, Xiaojie Zhou and Yueqiang Li
Buildings 2024, 14(10), 3119; https://doi.org/10.3390/buildings14103119 - 28 Sep 2024
Viewed by 526
Abstract
The horizontal joint is a critical component of the prefabricated shear wall structure, responsible for supporting both horizontal shear forces and vertical loads along with the wall, thereby influencing the overall structural performance. This study employs direct shear testing and finite element analysis [...] Read more.
The horizontal joint is a critical component of the prefabricated shear wall structure, responsible for supporting both horizontal shear forces and vertical loads along with the wall, thereby influencing the overall structural performance. This study employs direct shear testing and finite element analysis to investigate the horizontal joint in walls with ring reinforcement. It examines the impact of various factors on joint shear performance, including the type of joint material, joint configuration, buckling length of ring reinforcement, strength of precast concrete, reinforcement ratio of ring reinforcement and dowel bars, and the effect of horizontal binding force. The findings indicate that the shear bearing capacity and stiffness of joints incorporating post-cast epoxy resin concrete and keyways are comparable or superior to those of integrally cast specimens. A larger buckling length in ring reinforcement may reduce shear strength, suggesting an optimal buckling length at approximately one-third of the joint width. As the strength of precast concrete increases, ductility decreases while bearing capacity increases, initially at an increasing rate that subsequently declines. Optimal results are achieved when the strength of precast concrete closely matches that of the post-cast epoxy concrete. Enhancing the reinforcement ratio of ring reinforcement improves shear capacity, but excessively high ratios significantly reduce ductility. It is recommended that the diameter of ring reinforcement be maintained between 10 mm and 12 mm, with a reinforcement ratio between 0.79% and 1.13%. Increasing horizontal restraint enhances stiffness and shear capacity but reduces ductility; thus, the axial compression ratio should not exceed 0.5. Full article
(This article belongs to the Special Issue Advances in Novel Precast Concrete Structures)
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Figure 1
<p>Technical diagram of the buckle connection of ring reinforcement.</p>
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<p>Size and reinforcement diagram of typical specimen.</p>
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<p>Size and reinforcement diagram of typical specimen.</p>
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<p>Test device diagram.</p>
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<p>Loading system.</p>
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<p>Displacement meter layout.</p>
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<p>Failure condition of each specimen.</p>
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<p>Load-displacement curves of ZHY specimen.</p>
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<p>Failure of each specimen.</p>
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<p>Load-displacement curves of ZXJ-2-60 and ZXJ-2-100 specimen.</p>
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<p>Schematic diagram of ZHY-1-60 joint structure.</p>
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<p>Damage condition of each specimen.</p>
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<p>Load-displacement curves of ZXJ-1-60 and ZHY-1-60 specimen.</p>
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<p>Comparison of failure conditions of each specimen.</p>
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<p>Load-displacement curve of an ordinary concrete specimen with a 60mm length of U-shaped steel bar.</p>
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<p>Failure condition of specimen ZZJ.</p>
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<p>Comparison of load-displacement curves of all specimens.</p>
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<p>Finite element model of specimen ZHY-1-60.</p>
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<p>Comparison of test and simulation results of failure morphology of specimens.</p>
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<p>Comparison of experimental and finite element simulation results of load-displacement curves of specimens.</p>
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<p>Stress distribution of concrete and steel bars under peak load.</p>
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<p>Stress distribution of concrete and steel bars under peak load.</p>
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<p>Influence of precast partial concrete strength on joint shear capacity.</p>
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<p>Load-displacement curves of specimens with different annulus diameters.</p>
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<p>Comparison of the peak load of each specimen under different annulus diameters.</p>
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<p>Load displacement curve of each specimen under different dowel bar diameters.</p>
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<p>Peak load of each specimen under different insert bar diameters.</p>
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<p>Load displacement curve of each specimen under different horizontal binding forces.</p>
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<p>Peak load under different levels of restraint.</p>
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