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29 pages, 34281 KiB  
Article
Bio-Inspired Thin-Walled Straight and Tapered Tubes with Variable Designs Subjected to Multiple Impact Angles for Building Constructions
by Quanjin Ma, Nor Hazwani Mohd Yusof, Santosh Kumar Sahu, Yiheng Song, Nabilah Afiqah Mohd Radzuan, Bo Sun, Ahmad Yunus Nasution, Alagesan Praveen Kumar and Mohd Ruzaimi Mat Rejab
Buildings 2025, 15(4), 620; https://doi.org/10.3390/buildings15040620 - 17 Feb 2025
Abstract
Thin-walled structures are extensively utilized in construction because of their lightweight nature and excellent energy absorption efficiency, especially under dynamic loads. Improving the energy-absorbing performance of thin-walled structures by inspiring natural multi-cell designs is a sufficient approach. This paper investigates the energy-absorbing characteristics [...] Read more.
Thin-walled structures are extensively utilized in construction because of their lightweight nature and excellent energy absorption efficiency, especially under dynamic loads. Improving the energy-absorbing performance of thin-walled structures by inspiring natural multi-cell designs is a sufficient approach. This paper investigates the energy-absorbing characteristics of variable novel cross-section designs of thin-walled structures subjected to oblique impact loading. Straight and tapered types with seven cross-sectional designs of novel thin-walled structures were studied. The nonlinear ABAQUS/Explicit software 6.13 version was implemented to analyze the crashworthiness behaviors for the proposed variable cross-section designs under different loading angles. The crushing behaviors of the proposed thin-walled structures were examined for various wall thicknesses of 0.5 mm, 1.5 mm, and 2.5 mm and impact loading angles of 0°, 15°, 30°, and 45°. It was determined that the energy-absorbing characteristics of novel thin-walled structures can be efficiently controlled by varying two geometries and seven cross-section designs. A multi-criteria decision-making method (MCDM) using a complex proportional assessment method (COPRAS) was performed to select the optimum thin-walled structures with cross-section designs. It was shown that a tapered square thin-walled structure with 2.5 mm thickness had the best crashworthiness performances with energy absorption (EA) of 11.01 kJ and specific energy absorption (SEA) of 20.32 kJ/kg under a 30° impact angle. Moreover, the results indicated that the EA of the thin-walled structure decreased with the increase in the impact loading angle. In addition, with the increase in the impact loading angle, the peak crushing force (PCF) decreased and reflected the reduction in energy absorbed at a larger angle. The MCDM method in conjunction with the COPRAS method is proposed; it provides valuable insights for safer and more resilient building construction. Full article
(This article belongs to the Special Issue Bionic Materials and Structures in Civil Engineering)
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Figure 1

Figure 1
<p>Thin-walled structure as tubular roof truss element in civil engineering under multi-angle impacts.</p>
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<p>Bio-inspired design concepts from the cactus for building constructions in civil engineering.</p>
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<p>Seven designs of the thin-walled structures were used in this study: (<b>a</b>) straight tube; (<b>b</b>) tapered tube.</p>
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<p>Seven designs of the thin-walled structures were used in this study: (<b>a</b>) straight tube; (<b>b</b>) tapered tube.</p>
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<p>Procedure of finite element modeling: (<b>a</b>) Example of impact angle and boundary conditions of tapered nonagon tubes; (<b>b</b>) mesh convergence sensitivity analysis of thin-walled tube in terms of CPU time and percentage of correlation; (<b>c</b>) oblique impact angles from thin-walled truss in civil engineering.</p>
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<p>Deformation modes of thin-walled straight tubes with 0.5 mm wall thickness under 0°, 15°, 30°, and 45° impact angles.</p>
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<p>Deformation modes of thin-walled straight tubes with 1.5 mm wall thickness under 0°, 15°, 30°, and 45° impact angles.</p>
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<p>Deformation modes of thin-walled straight tubes with 2.5 mm wall thickness under 0°, 15°, 30°, and 45° impact angles.</p>
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<p>Deformation modes of thin-walled tapered tubes with 0.5 mm wall thickness under 0°, 15°, 30°, and 45° impact angles.</p>
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<p>Deformation modes of thin-walled tapered tubes with 1.5 mm wall thickness under 0°, 15°, 30°, and 45° impact angles.</p>
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<p>Deformation modes of thin-walled tapered tubes with 2.5 mm wall thickness under 0°, 15°, 30°, and 45° impact angles.</p>
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<p>Load–displacement curves of thin-walled straight tubes with 0.5 mm wall thickness subjected to multiple loading angles: (<b>a</b>) 0°; (<b>b</b>) 15°; (<b>c</b>) 30°; (<b>d</b>) 45°.</p>
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<p>Load–displacement curves of thin-walled straight tubes with 1.5 mm wall thickness subjected to multiple loading angles: (<b>a</b>) 0°; (<b>b</b>) 15°; (<b>c</b>) 30°; (<b>d</b>) 45°.</p>
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<p>Load–displacement curves of thin-walled straight tubes with 2.5 mm wall thickness subjected to multiple loading angles: (<b>a</b>) 0°; (<b>b</b>) 15°; (<b>c</b>) 30°; (<b>d</b>) 45°.</p>
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<p>Effect of impact angle and wall thickness of specific energy absorption (SEA) on thin-walled straight tubes under three wall thicknesses: (<b>a</b>) 0.5 mm; (<b>b</b>) 1.5 mm; (<b>c</b>) 2.5 mm.</p>
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<p>Effect of impact angle and wall thickness of specific energy absorption (SEA) on thin-walled tapered tubes under three wall thicknesses: (<b>a</b>) 0.5 mm; (<b>b</b>) 1.5 mm; (<b>c</b>) 2.5 mm.</p>
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<p>Results of SEA of thin-walled tubes as a function of impact angle and wall thickness: (<b>a</b>) straight type; (<b>b</b>) tapered type.</p>
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<p>Result of overall SEAα with thin-walled straight and tapered tubes with seven geometry profiles: (<b>a</b>) CASE I; (<b>b</b>) CASE II.</p>
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<p>Effect of impact angle and wall thickness of PCF on thin-walled straight tubes with three wall thicknesses: (<b>a</b>) 0.5 mm; (<b>b</b>) 1.5 mm; (<b>c</b>) 2.5 mm.</p>
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<p>Effect of impact angle and wall thickness of PCF on thin-walled tapered tubes with three wall thicknesses: (<b>a</b>) 0.5 mm; (<b>b</b>) 1.5 mm; (<b>c</b>) 2.5 mm.</p>
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<p>Results of COPRAS method for MCDM process on straight and tapered tubes with different designs: (<b>a</b>) optimum design; (<b>b</b>) worst design.</p>
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<p>Thin-walled structures used in civil engineering applications: (<b>a</b>) truss bridge; (<b>b</b>) roof truss structural framework; (<b>c</b>) transmission tower; (<b>d</b>) “Eye of Shenzhen” of Gangxia North Hub Station.</p>
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24 pages, 4137 KiB  
Article
Seismic Behavior of Composite Beam to Concrete-Filled Cold-Formed High-Strength Square Steel Tubular Column Joints with Different Connection Forms
by Jiangran Guo, Longhui Sun, He Zhao and Xihan Hu
Buildings 2025, 15(4), 622; https://doi.org/10.3390/buildings15040622 - 17 Feb 2025
Abstract
To enhance the standardization and construction efficiency of prefabricated steel structures and to promote the application of cold-formed steel tubes with the advantages of high standardization, superior mechanical properties, and fast processing speeds, two types of composite beam to concrete-filled cold-formed high-strength square [...] Read more.
To enhance the standardization and construction efficiency of prefabricated steel structures and to promote the application of cold-formed steel tubes with the advantages of high standardization, superior mechanical properties, and fast processing speeds, two types of composite beam to concrete-filled cold-formed high-strength square steel tubular column joints with different connection forms were designed in this study: the external diaphragm joint (ED joint) and the through diaphragm joint (TD joint). These joints were subjected to cyclic loading tests to evaluate the influence of the connection designs on key seismic performance parameters, such as failure modes, load-bearing capacities, the degradation of strength and stiffness, ductility, and energy dissipation capabilities. The results show that both the ED and TD joints experienced butt weld fractures at the bolted-welded connections on the beam, effectively transferring the plastic hinges from the joint zone to the beam and demonstrating good seismic performance. The ED joint specimen JD1 and the TD joint specimen JD2 exhibited similar load-bearing capacity, stiffness, strength degradation, and energy dissipation capacity. However, the TD joint showed lower ductility compared to the ED joint due to premature weld fractures. A nonlinear finite element model (FEM) was developed using MSC.MARC 2012, and the numerical simulation showed that the FEM could effectively simulate the hysteresis performance of the composite beam to concrete-filled, cold-formed, high-strength, square, steel tubular column joints with external and through diaphragms. Full article
(This article belongs to the Special Issue Advances in Structural Techniques for Prefabricated Modular Buildings)
21 pages, 13440 KiB  
Article
Dynamic Adaptability of Spherical Bearings in Small-Span Bridges for Heavy-Haul Railways
by Shuli Chen, Ye Zhou, Kaize Xie, Panhui Zhang and Chen Li
Buildings 2025, 15(4), 619; https://doi.org/10.3390/buildings15040619 - 17 Feb 2025
Abstract
Plate bearings in existing small-span bridges for heavy-haul railways have exhibited corrosion, detachment, and surface cracks under large axle loads, making them inadequate for the “capacity expansion and renovation” of heavy-haul railways. Therefore, identifying new bearings suitable for small-span bridges and developing a [...] Read more.
Plate bearings in existing small-span bridges for heavy-haul railways have exhibited corrosion, detachment, and surface cracks under large axle loads, making them inadequate for the “capacity expansion and renovation” of heavy-haul railways. Therefore, identifying new bearings suitable for small-span bridges and developing a rapid bearing replacement method tailored to the operational needs of heavy-haul railways are urgent priorities. This paper takes spherical bearings as an example and proposes a method for rapidly replacing plate bearings with spherical bearings. The bearing replacement tests of six simply supported beams were carried out to verify the effectiveness of the proposed method. Dynamic performance tests of bridges and bearings were performed before and after the replacement. A finite element model was established to analyze the effects of bridge span and pier height. The results show that the entire bearing replacement process for a span bridge could be completed within 4 h using the proposed method. Compared to plate bearings, spherical bearings could improve the lateral dynamic performance of both the bridge and bearings. However, the improvement decreases as bridge span and pier height increase. For 2.2 m diameter cylindrical piers commonly used in heavy-haul railways, the pier height with spherical bearings should be limited to 10 m. Full article
(This article belongs to the Topic Advances on Structural Engineering, 3rd Edition)
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Figure 1
<p>Various failures of plate bearings.</p>
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<p>Structural composition of spherical bearing.</p>
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<p>Beam lifting operation.</p>
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<p>Bearing replacement construction.</p>
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<p>Flowchart of bridge bearing replacement construction.</p>
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<p>Dimensions of bridge piers (Unit: m).</p>
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<p>Test bridge measuring point layout diagram.</p>
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<p>Field records.</p>
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<p>Dynamic deflection of the middle section.</p>
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<p>Maximum dynamic deflection.</p>
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<p>Lateral amplitude.</p>
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<p>Lateral acceleration.</p>
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<p>Maximum lateral amplitude.</p>
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<p>Maximum lateral acceleration.</p>
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<p>Maximum vertical amplitude.</p>
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<p>Maximum vertical acceleration.</p>
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<p>Lateral displacement at bearings.</p>
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<p>Vertical displacement at bearings.</p>
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<p>Lateral displacement amplitude.</p>
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<p>Vertical displacement amplitude.</p>
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<p>Maximum lateral amplitude at the pier top.</p>
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<p>Simulation model and beam cross-section.</p>
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<p>Vertical loads (Unit:m).</p>
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<p>Lateral forces.</p>
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<p>Dynamic response at a speed of 55 km/h (pier height).</p>
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<p>Maximum dynamic response (pier height).</p>
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<p>Maximum dynamic response (bridge span).</p>
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21 pages, 33783 KiB  
Article
Numerical Simulation of the Gas Flow of Combustion Products from Ignition in a Solid Rocket Motor Under Conditions of Propellant Creep
by Yin Zhang, Zhensheng Sun, Yu Hu, Yujie Zhu, Xuefeng Xia, Huang Qu and Bo Tian
Aerospace 2025, 12(2), 153; https://doi.org/10.3390/aerospace12020153 - 17 Feb 2025
Abstract
The development of modern solid rocket technology with high-performance and high-loading ratio propellants places higher requirements on the safety and stability of the solid rocket motor. The propellant of the solid rocket motor will creep during long-term vertical storage, which may adversely affect [...] Read more.
The development of modern solid rocket technology with high-performance and high-loading ratio propellants places higher requirements on the safety and stability of the solid rocket motor. The propellant of the solid rocket motor will creep during long-term vertical storage, which may adversely affect its regular operation. The ignition transient process is a critical phase in the operation of solid rocket motors. The Abaqusv.2022 finite element simulation software is used to analyze the ignition transient under propellant creep conditions and obtain the deformed combustion chamber profile. Then, we use a high-precision finite volume solver developed independently to simulate the flow field during the ignition process. In the simulation, we adopt the surface temperature of the propellant column reaching the ignition threshold as the ignition criterion, considering the heat transfer process of the propellant column instead of using the near-wall gas temperature to obtain the set temperature. Simulation results under different creep conditions reveal that the deformation of the propellant grains progressively intensifies as the solid rocket motor’s storage duration increases. This leads to a delayed initial ignition time of the propellant, an advancement of the overall ignition transient process, and an increased pressurization rate during ignition, which can affect the structure and regular operation of the motor. The research results provide design guidance and theoretical support for the design and life prediction of solid rocket motors. Full article
(This article belongs to the Section Astronautics & Space Science)
19 pages, 3933 KiB  
Article
A Fully Coupled Electro-Vibro-Acoustic Benchmark Model for Evaluation of Self-Adaptive Control Strategies
by Thomas Kletschkowski
J 2025, 8(1), 6; https://doi.org/10.3390/j8010006 - 17 Feb 2025
Viewed by 6
Abstract
The reduction of noise and vibration is possible with passive, semi-active and active control strategies. Especially where self-adaptive control is required, it is necessary to evaluate the noise reduction potential before the control approach is applied to the real-world problem. This evaluation can [...] Read more.
The reduction of noise and vibration is possible with passive, semi-active and active control strategies. Especially where self-adaptive control is required, it is necessary to evaluate the noise reduction potential before the control approach is applied to the real-world problem. This evaluation can be based on a virtual model that contains all relevant sub-systems, transfer paths and coupling effects on the one hand. On the other hand, the complexity of such a model has to be limited to focus on principal findings such as convergence speed, power consumption, and noise reduction potential. The present paper proposes a fully coupled electro-vibro-acoustic model for the evaluation of self-adaptive control strategies. This model consists of discrete electrical and mechanical networks that are applied to model the electro-acoustic behavior of noise and anti-noise sources. The acoustic field inside a duct, terminated by these electro-acoustic sources, is described by finite elements. The resulting multi-physical model is capable of describing all relevant coupling effects and enables an efficient evaluation of different control strategies such as the local control of sound pressure or active control of acoustic absorption. It is designed as a benchmark model for the benefit of the scientific community. Full article
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Figure 1
<p>Topological model of system (top) and electro-vibro-acoustical model (bottom).</p>
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<p>Resonance frequencies of the uncontrolled system.</p>
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<p>Normalized mode shapes in resonance.</p>
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<p>System input and system output without self-adaptive control.</p>
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<p>IR and resonance frequencies of the uncontrolled system.</p>
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<p>Modelling of system response without active control.</p>
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<p>Active control of local sound pressure—time-history of simulation.</p>
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<p>Frequency domain illustration of active control of local sound pressure.</p>
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<p>Active control of local absorption—time-history of simulation.</p>
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<p>Frequency domain illustration of active control of local absorption.</p>
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21 pages, 734 KiB  
Article
Nusselt Number Dependence on Friction Factor in the Boundary Slip Flow of a Newtonian Liquid Between Parallel Plates
by Krishna Kota, Sarada Kuravi and Prasanna Jayaramu
Thermo 2025, 5(1), 7; https://doi.org/10.3390/thermo5010007 - 17 Feb 2025
Viewed by 55
Abstract
This study explored the relationship between the Nusselt number and the friction factor in the laminar boundary slip flow of a Newtonian liquid between parallel plates. In addition, simplified equations were developed to estimate two key parameters—slip velocity and temperature jump—both of which [...] Read more.
This study explored the relationship between the Nusselt number and the friction factor in the laminar boundary slip flow of a Newtonian liquid between parallel plates. In addition, simplified equations were developed to estimate two key parameters—slip velocity and temperature jump—both of which are typically difficult to measure in experimental settings. The primary objectives of investigating the relationship between the Nusselt number and the friction factor were twofold: (1) to uncover the previously unknown mathematical connection (or analogy) between momentum transfer and heat transfer in the presence of boundary slip and (2) to enable predictions of either the pressure drop or the heat transfer coefficient by measuring just one of these quantities, thus simplifying experimental procedures. Considering the difficulty of conducting experiments of this type of flow (as described in the published literature), a finite element-based numerical model built in COMSOL Multiphysics software was used to validate the theoretically developed relationship over a wide range of Reynolds numbers and boundary slip values. While surface modifications like dimples, bumps, and ribs typically modify both the Nusselt number and pressure drop, leading to their increase for a given fluid and constant inlet Reynolds number, their behavior changes when boundary slip is present, particularly in cases where there is a low temperature jump at the wall. The analysis identified a specific threshold for the dimensionless temperature jump below which the Nusselt number with boundary slip will exceed 8.235. Furthermore, the analysis showed that for the Nusselt number to rise above 8.235, the non-dimensional velocity slip must be at least 3.19 times larger than the non-dimensional temperature jump. This means that the velocity slip has to be significantly larger than the temperature jump to achieve enhanced heat transfer in boundary slip flows. Full article
21 pages, 2027 KiB  
Article
Mathematical Modeling and Electromagnetic Characteristics Analysis of a Six-Phase Distributed Single-Winding BPMSM with 12 Slots and 2 Poles
by Wenshao Bu, Jiangdi Li and Yongfang Lu
Appl. Sci. 2025, 15(4), 2093; https://doi.org/10.3390/app15042093 - 17 Feb 2025
Viewed by 44
Abstract
This work focuses on small bearingless permanent magnet synchronous motors (BPMSMs). In order to enhance its torque control stiffness and improve the stability of its torque and magnetic levitation force dynamic waveforms, a novel six-phase distributed single-winding BPMSM with 12 slots and 2 [...] Read more.
This work focuses on small bearingless permanent magnet synchronous motors (BPMSMs). In order to enhance its torque control stiffness and improve the stability of its torque and magnetic levitation force dynamic waveforms, a novel six-phase distributed single-winding BPMSM with 12 slots and 2 poles (six-phase DSW-12/2-BPMSM) is proposed and researched in this work. First, the structure and working principle of the six-phase DSW-12/2-BPMSM are analyzed. Subsequently, considering the relative permeability of permanent magnets, mathematical models of the inductance matrix, electromagnetic torque and radial magnetic levitation force are established. Then, using the finite element method (FEM), the control characteristics of the electromagnetic torque and magnetic levitation force of the six-phase DSW-12/2-BPMSM are analyzed, and the mathematical model is verified. Finally, FEM simulation analysis and comparisons are conducted with a commonly used six-phase centralized single-winding BPMSM with 6 slots and 2 poles (six-phase CSW-6/2-BPMSM). The research results show that the established mathematical model is effective and accurate compared with the six-phase CSW-6/2-BPMSM. The six-phase DSW-12/2-BPMSM has greater torque control stiffness, its dynamic waveforms of torque and radial magnetic levitation force have higher quality and stability, and the coupling degree between its torque and radial magnetic levitation force is lower. Full article
(This article belongs to the Special Issue Power Electronics and Motor Control)
27 pages, 15329 KiB  
Review
Research Status and Development Trends of Joining Technologies for Ceramic Matrix Composites
by Biao Chen, Hang Sun, Yuchen Ye, Chunming Ji, Shidong Pan and Bing Wang
Materials 2025, 18(4), 871; https://doi.org/10.3390/ma18040871 - 17 Feb 2025
Viewed by 15
Abstract
Ceramic matrix composites (CMCs) are composite materials made by using structural ceramics as matrix and reinforcing components such as high-strength fibers, whiskers, or particles. These materials are combined in a specific way to achieve a composite structure. With their excellent properties, including high [...] Read more.
Ceramic matrix composites (CMCs) are composite materials made by using structural ceramics as matrix and reinforcing components such as high-strength fibers, whiskers, or particles. These materials are combined in a specific way to achieve a composite structure. With their excellent properties, including high specific strength, high specific stiffness, good thermal stability, oxidation resistance, and corrosion resistance, CMCs are widely used in the aerospace, automotive, energy, defense, and bio-medical fields. However, large and complex-shaped ceramic matrix composite parts are greatly influenced by factors such as the molding process, preparation costs, and consistency of quality, which makes the joining technology for CMCs increasingly important and a key trend for future development. However, due to the anisotropic nature of CMCs, the design of structural components varies, with different properties in different directions. Additionally, the chemical compatibility and physical matching between dissimilar materials in the joining process lead to much more complex joint design and strength analysis compared to traditional materials. This paper categorizes the joining technologies for CMCs into mechanical joining, bonding, soldering joining, and hybrid joining. Based on different joining techniques, the latest research progress on the joining of CMCs with themselves or with metals is reviewed. The advantages and disadvantages of each joining technology are summarized, and the future development trends of these joining technologies are analyzed. Predicting the performance of joining structures is currently a hot topic and challenge in research. Therefore, the study systematically reviews research combining failure mechanisms of ceramic matrix composite joining structures with finite element simulation techniques. Finally, the paper highlights the breakthroughs achieved in current research, as well as existing challenges, and outlines future research and application directions for ceramic matrix composite joining. Full article
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Figure 1
<p>Application of CMCs in various fields [<a href="#B1-materials-18-00871" class="html-bibr">1</a>].</p>
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<p>Finite element model and contact pair setup of 2D C/SiC composite—high-temperature alloy bolt joint structure [<a href="#B13-materials-18-00871" class="html-bibr">13</a>].</p>
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<p>Failure modes of C/SiC riveted joints: (<b>a</b>) Specimen with rivets; (<b>b</b>) specimen with rivets fully ejected [<a href="#B15-materials-18-00871" class="html-bibr">15</a>].</p>
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<p>Preparation of a two-dimensional C/SiC composite Z-direction pin joint with four rectangular arrays [<a href="#B19-materials-18-00871" class="html-bibr">19</a>].</p>
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<p>Preparation process of laminated and core–shell C/SiC bolts [<a href="#B20-materials-18-00871" class="html-bibr">20</a>].</p>
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<p>(<b>a</b>) Scanning electron microscope (SEM) cross-sectional micrograph of the CMC/EN AW—6082 joint using SUP42 HT-2 (Master Bond, The USA) adhesive and post-heat-treated at 200 °C for 10 min, and (<b>b</b>) magnified view of the interface [<a href="#B26-materials-18-00871" class="html-bibr">26</a>].</p>
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<p>SEM images of the joint and GOX microcrystalline glass crystallization. (<b>a</b>) Cross-sectional view of the joint; (<b>b</b>) cross-sectional view of the joint in backscattered electron mode (the arrow indicates the bonding interface); (<b>c</b>) crystalline phase (bright color) in the glass matrix (dark color); (<b>d</b>) cross-sectional view of the joint after thermal aging for 100 h at 850 °C in air (the arrow indicates the bonding interface) [<a href="#B28-materials-18-00871" class="html-bibr">28</a>].</p>
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<p>Optical micrographs of the vertical SA-Tyrannohex assembly (⊥-to-⊥) bonded using (<b>a</b>) Si-Cr, (<b>b</b>) Si-Ti, and (<b>c</b>) Si-Hf eutectic phase pastes [<a href="#B32-materials-18-00871" class="html-bibr">32</a>].</p>
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<p>Microstructural evolution mechanisms: (<b>a</b>) before joining; (<b>b</b>) during brazing, joint based on Ticusil filler; (<b>c</b>) after brazing, joint based on Ticusil filler; (<b>d</b>) after brazing, joint based on Cusil filler; (<b>e</b>) during LSS testing, joint based on Ti-Cusil filler; (<b>f</b>) during LSS testing, joint based on Cusil filler [<a href="#B36-materials-18-00871" class="html-bibr">36</a>].</p>
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<p>BSE images of the joint interface microstructure brazed for 10 min at different temperatures [<a href="#B39-materials-18-00871" class="html-bibr">39</a>].</p>
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<p>Diagram of Ti<sub>16</sub>Si<sub>84</sub> alloy joining SiC<sub>f</sub>/SiC composite [<a href="#B42-materials-18-00871" class="html-bibr">42</a>].</p>
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<p>(<b>a</b>) XRD patterns of the C/SiC fracture after single-lap shear strength test; (<b>b</b>) Shear strength of C/SiC joint; 3D images and BSE images of C/SiC joint fracture: (<b>c</b>) NS1 1000 °C, (<b>d</b>) NS2 1200 °C, (<b>e</b>) NS3 1400 °C, and (<b>f</b>) NS4 1600 °C [<a href="#B43-materials-18-00871" class="html-bibr">43</a>].</p>
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<p>Explanation of the preparation and assembly process of the hybrid joint: (<b>a</b>) material orientation of the substrate and screw; (<b>b</b>) assembly process of the hybrid joint [<a href="#B48-materials-18-00871" class="html-bibr">48</a>].</p>
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<p>Three types of SiC deposition zones in the C/SiC z-pinned/bonded hybrid single-lap joint: (<b>a</b>) deposition zone A: near the edge; (<b>b</b>) deposition zone B: near the z-pin; (<b>c</b>) deposition zone C: gap between the z-pin and hole; (<b>d</b>) numerical model; (<b>e</b>) distribution of fasteners with similar area in deposition zone A [<a href="#B49-materials-18-00871" class="html-bibr">49</a>].</p>
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<p>Failure comparison between numerical simulation and experimental results: (<b>a</b>) numerical simulation failure mode and predicted progressive failure process; (<b>b</b>) numerical simulation failure diagram of the bolt; (<b>c</b>) typical failure modes in the experiment [<a href="#B58-materials-18-00871" class="html-bibr">58</a>].</p>
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<p>Simulation of the failure process: (<b>a</b>) SiC bonding layer in the semi-lap region; (<b>b</b>) SiC bonding layer in the entire lap region [<a href="#B60-materials-18-00871" class="html-bibr">60</a>].</p>
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26 pages, 6335 KiB  
Article
Analysis of Nonlinear Dynamics of a Gear Transmission System Considering Effects of the Extended Tooth Contact
by Fulin Liao, Xingyuan Zheng, Jianliang Huang and Weidong Zhu
Machines 2025, 13(2), 155; https://doi.org/10.3390/machines13020155 - 17 Feb 2025
Viewed by 8
Abstract
Considering the elasticity of gear solid bodies, the load applied to gear teeth will force theoretically separated gear teeth to get into engaging state in advance. This phenomenon is named as the extended tooth contact (ETC). Effects of the ETC directly influence the [...] Read more.
Considering the elasticity of gear solid bodies, the load applied to gear teeth will force theoretically separated gear teeth to get into engaging state in advance. This phenomenon is named as the extended tooth contact (ETC). Effects of the ETC directly influence the time-varying mesh stiffness of gear pairs and subsequently alter nonlinear dynamic characteristics of gear transmission systems. Time-vary mesh stiffness, considering effects of the ETC, is thus introduced into the dynamic model of the gear transmission system. Periodic motions of a gear transmission system are discussed in detail in this work. The analytical model of time-varying mesh stiffness with effects of the ETC is proposed, and the effectiveness of the analytical model is demonstrated in comparison with finite element (FE) results. The gear transmission system is simplified as a single degree-of-freedom (DOF) model system by employing the lumped mass method. The correctness of the dynamic model is verified in comparison with experimental results. An incremental harmonic balance (IHB) method is modified to obtain periodic responses of the gear transmission system. The improved Floquet theory is employed to determine the stability and bifurcation of the periodic responses of the gear transmission system. Some interesting phenomena exist in the periodic responses consisting of “softening-spring” behaviors, jump phenomena, primary resonances (PRs), and super-harmonic resonances (SP-HRs), and saddle-node bifurcations are observed. Especially, effects of loads on unstable regions, amplitudes, and positions of bifurcation points of frequency response curves are revealed. Analytical results obtained by the IHB method match very well with those from numerical integration. Full article
(This article belongs to the Special Issue Advancements in Mechanical Power Transmission and Its Elements)
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<p>The single-DOF dynamic model of the gear transmission system considering time-varying mesh stiffness.</p>
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<p>A logic diagram the IHB method.</p>
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<p>Comparisons between analytical results and FE results of the time-varying mesh stiffness of the gear transmission system for loads of 100 Nm and 180 Nm: (<b>a</b>) the gear transmission system neglecting effects of the ETC for a load of 100 Nm, (<b>b</b>) the gear transmission system considering effects of the ETC for a load of 100 Nm, (<b>c</b>) the gear transmission system neglecting effects of the ETC for a load of 180 Nm, and (<b>d</b>) the gear transmission system considering effects of the ETC for a load of 180 Nm.</p>
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<p>Comparisons between analytical results and FE results of the time-varying mesh stiffness of the gear transmission system for loads of 100 Nm and 180 Nm: (<b>a</b>) the gear transmission system neglecting effects of the ETC for a load of 100 Nm, (<b>b</b>) the gear transmission system considering effects of the ETC for a load of 100 Nm, (<b>c</b>) the gear transmission system neglecting effects of the ETC for a load of 180 Nm, and (<b>d</b>) the gear transmission system considering effects of the ETC for a load of 180 Nm.</p>
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<p>Effects of loads on the time-varying mesh stiffness of the gear transmission system: (<b>a</b>) the gear transmission system neglecting effects of the ETC and (<b>b</b>) the gear transmission system considering effects of the ETC.</p>
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<p>Flowchart of collecting and analyzing experimental data.</p>
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<p>Equivalent root-mean-square (rms) amplitude <math display="inline"><semantics> <msub> <mi>A</mi> <mrow> <mi>r</mi> <mi>m</mi> <mi>s</mi> </mrow> </msub> </semantics></math> from experimental results [<a href="#B49-machines-13-00155" class="html-bibr">49</a>], numerical results [<a href="#B35-machines-13-00155" class="html-bibr">35</a>], and the IHB method results.</p>
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<p>Time history map and phase plane diagram of the periodic response when <math display="inline"><semantics> <mrow> <mi>ω</mi> <mo>=</mo> <mn>0.5990</mn> </mrow> </semantics></math>: (<b>a</b>) time history map and (<b>b</b>) phase plane diagram.</p>
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<p>Nonlinear dynamic responses of the gear transmission system under different loads: (<b>a</b>) the equivalent root-mean-square amplitude <math display="inline"><semantics> <msub> <mi>A</mi> <mrow> <mi>r</mi> <mi>m</mi> <mi>s</mi> </mrow> </msub> </semantics></math> of the gear transmission system neglecting effects of the ETC and (<b>b</b>) the equivalent root-mean-square amplitude <math display="inline"><semantics> <msub> <mi>A</mi> <mrow> <mi>r</mi> <mi>m</mi> <mi>s</mi> </mrow> </msub> </semantics></math> of the gear transmission system considering effects of the ETC; (<b>c</b>) the first-order harmonic amplitude <math display="inline"><semantics> <msub> <mi>A</mi> <mn>1</mn> </msub> </semantics></math> of the gear transmission system neglecting effects of the ETC and (<b>d</b>) the first-order harmonic amplitude <math display="inline"><semantics> <msub> <mi>A</mi> <mn>1</mn> </msub> </semantics></math> of the gear transmission system considering effects of the ETC; and (<b>e</b>) the second-order harmonic amplitude <math display="inline"><semantics> <msub> <mi>A</mi> <mn>2</mn> </msub> </semantics></math> of the gear transmission system neglecting effects of the ETC and (<b>f</b>) the second-order harmonic amplitude <math display="inline"><semantics> <msub> <mi>A</mi> <mn>2</mn> </msub> </semantics></math> of the gear transmission system considering effects of the ETC.</p>
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<p>Effects of loads on the bifurcation characteristics of the gear transmission system: (<b>a</b>) bifurcation characteristics of the gear transmission system neglecting the effects of the ETC and (<b>b</b>) bifurcation characteristics of the gear transmission system considering the effects of the ETC.</p>
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<p>Zeros of the equation <math display="inline"><semantics> <mrow> <mrow> <mo>|</mo> </mrow> <msub> <mover accent="true"> <mi>x</mi> <mo>¯</mo> </mover> <mn>0</mn> </msub> <mrow> <mo>|</mo> <mo>=</mo> <mi>b</mi> <mo>/</mo> </mrow> <msub> <mi>b</mi> <mi>c</mi> </msub> </mrow> </semantics></math>.</p>
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26 pages, 10090 KiB  
Article
Wear Resistance of Additively Manufactured Footwear Soles
by Shuo Xu, Shuvodeep De, Meysam Khaleghian and Anahita Emami
Lubricants 2025, 13(2), 89; https://doi.org/10.3390/lubricants13020089 - 17 Feb 2025
Viewed by 9
Abstract
This study systematically evaluated the wear resistance and mechanical performance of 3D-printed thermoplastic rubber (TPR) and flexible stereolithography (SLA) resin materials for footwear outsoles. Abrasion tests were conducted on 26 samples (2 materials × 13 geometries) to analyze the weight loss, variations in [...] Read more.
This study systematically evaluated the wear resistance and mechanical performance of 3D-printed thermoplastic rubber (TPR) and flexible stereolithography (SLA) resin materials for footwear outsoles. Abrasion tests were conducted on 26 samples (2 materials × 13 geometries) to analyze the weight loss, variations in the friction coefficient, temperature change, and deformation behavior. Finite element method (FEM) simulations incorporating the Ogden hyperelastic model were employed to investigate the stress distribution and wear patterns. The results revealed that TPR exhibits superior abrasion resistance and stable wear curves, making it suitable for high-load applications. On average, the TPR samples showed 27.3% lower weight loss compared to the SLA resin samples. The SLA resin samples exhibited a 65% higher mean coefficient of friction (COF) compared to the TPR samples. Furthermore, the SLA resin samples demonstrated a 94% higher temperature change during the sliding tests, reflecting greater friction-induced heating. The FEM simulations further validated TPR’s performance in high-stress regions and SLA resin’s deformation characteristics. This study’s findings not only highlight the performance differences between these two 3D-printed materials but also provide theoretical guidance for material selection based on wear behavior, contributing to the optimization of outsole design and its practical applications. Full article
(This article belongs to the Special Issue Wear and Friction in Hybrid and Additive Manufacturing Processes)
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<p>Pre-printing sample models for the Atomstack Cambrian 3D printer (1–13) and SLA printer (14–26) and the printed samples before abrasion testing.</p>
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<p>Sliding friction and wear test setup.</p>
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<p>The abrasion test friction force for the SLA resin sample 19 and TPR sample 6.</p>
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<p>The abrasion test weight loss of the testing samples: (<b>a</b>) TPR, and (<b>b</b>) SLA resin sample comparison.</p>
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<p>Contrast between the weight loss of sample 4 (TPR) and sample 17 (Resin) over the distances of testing.</p>
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<p>The mean COFs of concrete conditions for samples 1 through 26 for the first test.</p>
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<p>Key factors influencing wear.</p>
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<p>The wear damage results after an abrasion test: (<b>a</b>,<b>b</b>) depict the TPR samples, and (<b>c</b>–<b>e</b>) depict the resin samples.</p>
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<p>(<b>a</b>) The temperature change of samples 1 through 26. (<b>b</b>) Comparison of the temperature change between TRR and SLA.</p>
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<p>The wear and degradation results of different material surfaces after the abrasion test.</p>
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<p>FEM results compared with the experimental deformation for sample 17.</p>
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<p>FEM results compared with the experimental deformation for samples 1 vs. 14, 2 vs. 15, and 6 vs. 19.</p>
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<p>FEM results compared with the experimental deformation for samples 1 and 14.</p>
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<p>FEM results compared with the experimental deformation for shape-shifting sample 10.</p>
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<p>Samples printed with resin material.</p>
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<p>Model and boundary condition.</p>
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<p>(<b>a</b>) Coarse mesh with an average element edge length of 1 mm. (<b>b</b>) Refined mesh with an average element edge length of 0.8 mm. (<b>c</b>) von Mises stress plot for the coarse mesh. (<b>d</b>) von Mises stress plot for the coarse mesh.</p>
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18 pages, 3816 KiB  
Article
Experimental Investigation and FEM Simulation of the Tensile Behavior of Hot-Rolled Quenching and Partitioning 5Mn Steel
by Firew Tullu Kassaye, Tamiru Hailu Kori, Aleksandra Kozłowska and Adam Grajcar
Materials 2025, 18(4), 868; https://doi.org/10.3390/ma18040868 - 17 Feb 2025
Viewed by 22
Abstract
Medium manganese steels provide a good combination of tensile strength and ductility due to their multiphase microstructure produced during the multi-step heat treatment process. This study primarily focused on testing and analyzing the tensile properties of 0.17C-5Mn-0.76Al-0.9Si-Nb medium manganese quenching and partitioning (QP) [...] Read more.
Medium manganese steels provide a good combination of tensile strength and ductility due to their multiphase microstructure produced during the multi-step heat treatment process. This study primarily focused on testing and analyzing the tensile properties of 0.17C-5Mn-0.76Al-0.9Si-Nb medium manganese quenching and partitioning (QP) steel using both the experimental and finite element method (FEM) in the multilinear isotropic hardening material model. The 7 mm and 12 mm thick plates exhibited a similar microstructure of tempered primary martensite, lath-type retained austenite, and secondary martensite. The experiments measured tensile strengths of 1400 MPa for 12 mm round specimens and 1325 MPa for 7 mm flat specimens, with total elongations of 15% for round specimens and 11% for flat specimens. The results indicated that the sample’s geometry has some effect on the UTS and ductility of the studied medium-Mn QP steel. However, the more important is the complex relationship between the plate thickness and yield stress and ductility, which are affected by finishing hot rolling conditions. The FEM results showed that the von Mises stresses for flat and round specimens were 1496 MPa and 1514 MPa, respectively, and were consistent with the calculated true stresses of experimental results. This shows that numerical modeling, specifically a multilinear isotropic hardening material model, properly describes the material properties beyond the yield stress and accurately predicts the plastic deformation of the investigated multiphase QP steel. Full article
(This article belongs to the Section Metals and Alloys)
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<p>Processing schedule of investigated steel (FRT—finishing rolling temperature).</p>
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<p>Tensile test specimen dimensions: (<b>a</b>) round specimen from 12 mm thickness plate; (<b>b</b>) flat specimen from 7 mm thickness plate.</p>
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<p>Tensile test specimen dimensions: (<b>a</b>) round specimen from 12 mm thickness plate; (<b>b</b>) flat specimen from 7 mm thickness plate.</p>
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<p>Meshing: (<b>a</b>) round specimen; (<b>b</b>) flat specimen; boundary condition: (<b>c</b>) flat specimen; (<b>d</b>) round specimen: the round specimen from 12 mm thickness plate; the flat specimen from 7 mm thickness plate.</p>
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<p>SEM micrographs after quenching at 240 °C and partitioning at 450 °C (300 s): (<b>a</b>) plate 7 mm; (<b>b</b>) plate 12 mm. PM—primary martensite; SM—secondary martensite; RA—retained austenite.</p>
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<p>Comparison of averaged engineering and true stress-strain curves of investigated Q240P450(300 s) steel: (<b>a</b>) 12 mm thickness round specimen; (<b>b</b>) 7 mm thickness flat specimen.</p>
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<p>Strain hardening rate curves as a function of true strain for both plate thickness.</p>
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<p>FEM simulation results of equivalent (von-Mises) stresses for Q240P450(300 s): (<b>a</b>) flat specimen; (<b>b</b>) round specimen.</p>
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<p>Comparison of experimental and FEM stress-strain curves: (<b>a</b>) Q240P450(300 s)—12 mm plate thickness; (<b>b</b>) Q240P450(300 s)—7 mm plate thickness.</p>
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25 pages, 15494 KiB  
Article
Multi-Objective Geometry Optimization of Additive-Manufactured Hexagonal Honeycomb Sandwich Beams Under Quasi-Static Three-Point Bending Loading
by Andres Cecchini, Marco Menegozzo and Emerson Roman
Materials 2025, 18(4), 867; https://doi.org/10.3390/ma18040867 - 17 Feb 2025
Viewed by 125
Abstract
This research paper presents the findings of a design optimization analysis conducted on additive-manufactured thermoplastic sandwich structures with hexagonal honeycombs subjected to quasi-static three-point bending. Based on experimental results, finite element analysis, and analytical models, the relationship between four selected design variables (i.e., [...] Read more.
This research paper presents the findings of a design optimization analysis conducted on additive-manufactured thermoplastic sandwich structures with hexagonal honeycombs subjected to quasi-static three-point bending. Based on experimental results, finite element analysis, and analytical models, the relationship between four selected design variables (i.e., cell wall length ratio, cell wall angle, cell wall thickness, and skin thickness) and the structure’s mass, flexural stiffness, and maximum load capacity was determined. The influence of each design variable on the aforementioned structural properties was mathematically represented using three scaling laws to formulate a multi-objective optimization problem. Two conflicting objective functions, one for the mass and the other for the reciprocal of the maximum load capacity, along with a nonlinear constraint equation for the minimum allowed flexural stiffness of the sandwich structure were developed. The optimal values of the design variables were determined using two optimization methods, the Pareto optimal front and genetic algorithm, and by applying the Improved Minimum Distance Selection Method (IMDSM). Optimized designs were obtained for different values of flexural stiffness. It was found that, independently of the stiffness constraint value, the optimal value of the cell wall length ratio was 0.2 and the optimal cell wall thickness was 1.4 mm, which correspond to the minimum cell wall length ratio and maximum cell wall thickness considered in this study, respectively. On the other hand, if higher flexural stiffness is required for the structure, both cell wall angle and skin thickness must be increased accordingly. Furthermore, an increase in flexural stiffness is accompanied by an increase in both the mass and maximum load capacity of the structure. Full article
(This article belongs to the Special Issue Lightweight and High-Strength Sandwich Panel)
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<p>Honeycomb cell geometry and sandwich beam design variables.</p>
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<p>The 3D printing of sandwich beam test specimens.</p>
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<p>Tension test of PLA dogbone specimens: stress–strain curve (with the elastic branch zoomed-in on) and experimental setup.</p>
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<p>Quasi-static three-point bending test of 3D printed specimen.</p>
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<p>FE model of quasi-static three-point bending (implicit) analysis.</p>
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<p>Quasi-static three-point bending test of 3D printed specimens (h/l = 0.3).</p>
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<p>Quasi-static three-point bending test force–displacement curves (h/l = 0.3).</p>
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<p>Quasi-static three-point bending test of 3D printed specimens (h/l = 0.5).</p>
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<p>Quasi-static three-point bending test force–displacement curves (h/l = 0.5).</p>
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<p>Specimen mass.</p>
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<p>Specimen stiffness.</p>
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<p>Specimens’ maximum initial load.</p>
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<p>Specimens’ specific stiffness.</p>
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<p>Experimental and numerical force–displacement curves (h/l = 0.3).</p>
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<p>Experimental and numerical force–displacement curves (h/l = 0.5).</p>
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<p>FE (bars) vs. experimental (error bars) stiffness.</p>
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<p>FE (bars) vs. experimental (error bars) maximum initial load.</p>
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<p>Formation of plastic hinge in inclined honeycomb cell wall below the loading head.</p>
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<p>Relationship between h/l and beam mass, flexural stiffness, and maximum initial load.</p>
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<p>Relationship between θ and beam mass, flexural stiffness, and maximum initial load.</p>
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<p>Relationship between t and beam mass, flexural stiffness, and maximum initial load.</p>
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<p>Relationship between t<sub>f</sub> and beam mass, flexural stiffness, and maximum initial load.</p>
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<p>Pareto optimal front and GA (<span class="html-italic">K</span><sub>min</sub> = 0).</p>
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<p>Pareto front and GA (<span class="html-italic">K</span><sub>min</sub> = 500 N/mm).</p>
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17 pages, 5232 KiB  
Article
Topology Optimization and Testing of Connecting Rod Based on Static and Dynamic Analyses
by Mahalingam Nainaragaram Ramasamy, Aleš Slíva, Prasath Govindaraj and Akash Nag
Appl. Sci. 2025, 15(4), 2081; https://doi.org/10.3390/app15042081 - 16 Feb 2025
Viewed by 327
Abstract
This research article outlines our aim to perform topology optimization (TO) by reducing the mass of the connecting rod of an internal combustion engine based on static structural and dynamic analyses. The basic components of an internal combustion engine like the connecting rods, [...] Read more.
This research article outlines our aim to perform topology optimization (TO) by reducing the mass of the connecting rod of an internal combustion engine based on static structural and dynamic analyses. The basic components of an internal combustion engine like the connecting rods, pistons, crankshaft, and cylinder liners were designed using Autodesk Inventor Professional 2025. Using topology optimization, we aimed to achieve lesser maximum von Mises stress during static structural analysis and maintain a factor of safety (FOS) above 2.5 during rigid body dynamics. A force of 64,500 N was applied at the small end of the connecting rod while the big end was fixed. Topology optimization was carried out using ANSYS Discovery software at various percentages on a trial-and-error basis to determine better topology with lesser maximum von Mises stress. Target reduction was set to 4%, and as a result, 5.66% mass reduction from the original design and 6.25% reduced maximum von Mises stress was achieved. Later, transient analysis was carried out to evaluate the irregular motion loads and moments acting on the connecting rod at 1000 rpm. The results showed that the FOS remained above 2.5. Finally, the optimized connecting rod was simulated and verified for longevity using Goodman fatigue life analysis. Full article
(This article belongs to the Special Issue Computer-Aided Design in Mechanical Engineering)
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<p>Major dimensions of the connecting rod’s initial design.</p>
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<p>Virtual assembly of the crankshaft, connecting rod, piston, and cylinder liner.</p>
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<p>Topology optimization and validation process workflow.</p>
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<p>(<b>a</b>) Boundary conditions applied during the static structural analysis in the initial design. (<b>b</b>) Mesh properties used during the static structural analysis in the initial design.</p>
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<p>Protected depth of 0.012 m on both ends of the connecting rod.</p>
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<p>The sequence of the connecting rod TO from initial facets to the final subD hybrid conversion.</p>
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<p>(<b>a</b>) Original shape of the connecting rod. (<b>b</b>) 5.66% reduced mass from the original shape of the connecting rod.</p>
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<p>(<b>a</b>) Boundary conditions—A rotational velocity of 104.72 rad/s was applied to the crankshaft. (<b>b</b>,<b>c</b>) Maximum unknown reaction forces extracted using a joint probe while the piston was moving toward the top dead center and the bottom dead center.</p>
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<p>(<b>a</b>) Maximum unknown reaction forces applied during static structural analysis on the initial design while at the TDC. (<b>b</b>) Maximum unknown reaction forces applied during static structural analysis on the optimized topology with a mass reduction of 5.66% while at the TDC.</p>
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<p>(<b>a</b>) Maximum unknown reaction forces applied during static structural analysis on the initial design while at the BDC. (<b>b</b>) Maximum unknown reaction forces applied during static structural analysis on the optimized topology with a mass reduction of 5.66% while at the BDC.</p>
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<p>(<b>a</b>) Maximum von Mises stress acting on the initial design. (<b>b</b>) Maximum von Mises stress acting on the 5.66% mass-reduced optimized topology.</p>
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<p>(<b>a</b>) Safety factor after applying the motion loads on the 5.66% mass-reduced optimized topology at the TDC. (<b>b</b>) Safety factor after applying the motion loads on the 5.66% mass-reduced optimized topology at the BDC.</p>
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<p>(<b>a</b>) Life of the connecting rod after applying the motion loads on the 5.66% mass-reduced optimized topology at the TDC. (<b>b</b>) Life of the connecting rod after applying the motion loads on the 5.66% mass-reduced optimized topology at the BDC.</p>
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15 pages, 2667 KiB  
Article
Modal Analysis and Optimization of Tractor Exhaust System
by Ayla Tekin and Halil Şamlı
Appl. Sci. 2025, 15(4), 2070; https://doi.org/10.3390/app15042070 - 16 Feb 2025
Viewed by 165
Abstract
Excessive vibrations in exhaust systems can significantly reduce a vehicle’s lifespan and compromise performance. These vibrations, caused by factors such as engine operation and road conditions, lead to wear and tear. To address this issue, a finite element analysis (FEA) was conducted on [...] Read more.
Excessive vibrations in exhaust systems can significantly reduce a vehicle’s lifespan and compromise performance. These vibrations, caused by factors such as engine operation and road conditions, lead to wear and tear. To address this issue, a finite element analysis (FEA) was conducted on a 90-horsepower tractor’s exhaust system. Using ANSYS WB®, a 3D model was created and modal analysis was performed to determine the system’s natural frequencies and mode shapes. Based on the results, geometric modifications were made to the exhaust system, increasing its stiffness and shifting vibration frequencies to higher values. Consequently, vibration levels, noise, and the risk of component failure were significantly reduced. The redesigned exhaust system was successfully implemented in production. This study demonstrates the effectiveness of FEA in analyzing exhaust system vibrations and facilitating design improvements. By extending vehicle lifespan and providing a quieter, more comfortable driving experience, this research offers valuable insights for automotive and mechanical engineers. Full article
(This article belongs to the Special Issue Design and Optimization of Manufacturing Systems, 2nd Edition)
14 pages, 8059 KiB  
Article
The Effect of Through-Silicon-Via Thermal Stress on Metal-Oxide-Semiconductor Field-Effect Transistor Properties Under Cooling to Ultra-Low Temperatures
by Wenting Xie, Xiaoting Chen, Liting Zhang, Xiangjun Lu, Bing Ding and An Xie
Micromachines 2025, 16(2), 221; https://doi.org/10.3390/mi16020221 - 15 Feb 2025
Viewed by 213
Abstract
The thermal through-silicon-via (TTSV) has a serious thermal stress problem due to the mismatch of the coefficient of thermal expansion between the Si substrate and filler metal. At present, the thermal stress characteristics and strain mechanism of TTSV are mainly concerned with increases [...] Read more.
The thermal through-silicon-via (TTSV) has a serious thermal stress problem due to the mismatch of the coefficient of thermal expansion between the Si substrate and filler metal. At present, the thermal stress characteristics and strain mechanism of TTSV are mainly concerned with increases in temperature, and its temperature range is concentrated between 173 and 573 K. By employing finite element analysis and a device simulation method based on temperature-dependent material properties, the impact of TTSV thermal stress on metal-oxide-semiconductor field-effect transistor (MOSFET) properties is investigated under cooling down from room temperature to the ultra-low temperature (20 mK), where the magnitude of thermal stress in TTSV is closely associated with the TTSV diameter and results in significant tension near the Cu-Si interface and consequently increasing the likelihood of delamination and cracking. Considering the piezoresistive effect of the Si substrate, both the TTSV diameter and the distance between TTSV and MOSFET are found to have more pronounced effects on electron mobility along [100] crystal orientation and hole mobility along [110] crystal orientation. Applying a gate voltage of 3 V, the saturation current for the 45 nm-NMOS transistor oriented along channel [100] experiences a variation as high as 34.3%. Moreover, the TTSV with a diameter of 25 μm generates a change in MOSFET threshold voltage up to −56.65 mV at a distance as short as 20 μm. The influences exerted by the diameter and distance are consistent across carrier mobility, saturation current, and threshold voltage parameters. Full article
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<p>The schematic of TTSV model.</p>
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<p>Modeling domain of the unit cell model.</p>
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<p>Cooling process of the TTSV model with the PPMS refrigeration system.</p>
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<p>Temperature-dependent material properties of Si and Cu: (<b>a</b>) CTE, (<b>b</b>) Young’s modulus, (<b>c</b>) Poisson’s ratio, and (<b>d</b>) ultimate strength.</p>
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<p>TTSV deformation induced by thermal stress when temperature-dependent material properties are (<b>a</b>) not considered and (<b>b</b>) considered.</p>
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<p>Thermal stress distribution in the TTSV and location of maximum stress: (<b>a</b>) von Mises, (<b>b</b>) radial, (<b>c</b>) axial, and (<b>d</b>) shear stress.</p>
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<p>(<b>a</b>) Von Mises, (<b>b</b>) radial, (<b>c</b>) axial, and (<b>d</b>) shear stress for various Dias and Asps.</p>
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<p>Schematic diagram of MOSFET distribution in the [100] (<b>a</b>) and [110] (<b>b</b>) channel directions around TTSV.</p>
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<p>Contour plots of (<b>a</b>,<b>c</b>) electron and (<b>b</b>,<b>d</b>) hole mobility variations along the (<b>a</b>,<b>b</b>) [100] and (<b>c</b>,<b>d</b>) [110] crystal orientations.</p>
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<p>Influence of (<b>a</b>) Dia, (<b>b</b>) Asp, and (<b>c</b>) Dis on variation rate of carrier mobility.</p>
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<p>Optimal device placement of (<b>a</b>) NMOS in the [100] channel direction and (<b>b</b>) PMOS in the [110] channel direction.</p>
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<p>Structural profiles of (<b>a</b>) NMOS and (<b>b</b>) PMOS.</p>
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<p>Characteristic curves with a Dia of 25 μm for (<b>a</b>,<b>b</b>) NMOS and (<b>c</b>,<b>d</b>) PMOS with (<b>a</b>,<b>c</b>) [100] and (<b>b</b>,<b>d</b>) [110] channel directions.</p>
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<p>Effect of the (<b>a</b>) Dia, (<b>b</b>) Asp, and (<b>c</b>) Dis on the variation rate of the saturation current.</p>
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<p>Functional relationship between the biaxial strain and threshold voltage.</p>
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<p>Threshold voltage variations under criteria (<b>a</b>) ① and (<b>b</b>) ②.</p>
Full article ">Figure 17
<p>Impact of (<b>a</b>) Dia and (<b>b</b>) Dis on threshold voltage drift.</p>
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