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13 pages, 11404 KiB  
Essay
The Tectonic Significance of the Mw7.1 Earthquake Source Model in Tibet in 2025 Constrained by InSAR Data
by Shuyuan Yu, Shubi Zhang, Jiaji Luo, Zhejun Li and Juan Ding
Remote Sens. 2025, 17(5), 936; https://doi.org/10.3390/rs17050936 - 6 Mar 2025
Abstract
On 7 January 2025, at Beijing time, an Mw7.1 earthquake occurred in Dingri County, Shigatse, Tibet. To accurately determine the fault that caused this earthquake and understand the source mechanism, this study utilized Differential Interferometric Synthetic Aperture Radar (DInSAR) technology to [...] Read more.
On 7 January 2025, at Beijing time, an Mw7.1 earthquake occurred in Dingri County, Shigatse, Tibet. To accurately determine the fault that caused this earthquake and understand the source mechanism, this study utilized Differential Interferometric Synthetic Aperture Radar (DInSAR) technology to process Sentinel-A data, obtaining the line-of-sight (LOS) co-seismic deformation field for this earthquake. This deformation field was used as constraint data to invert the geometric parameters and slip distribution of the fault. The co-seismic deformation field indicates that the main characteristics of the earthquake-affected area are vertical deformation and east-west extension, with maximum deformation amounts of 1.6 m and 1.0 m for the ascending and descending tracks, respectively. A Bayesian method based on sequential Monte Carlo sampling was employed to invert the position and geometric parameters of the fault, and on this basis, the slip distribution was inverted using the steepest descent method. The inversion results show that the fault has a strike of 189.2°, a dip angle of 40.6°, and is classified as a westward-dipping normal fault, with a rupture length of 20 km, a maximum slip of approximately 4.6 m, and an average slip angle of about −82.81°. This indicates that the earthquake predominantly involved normal faulting with a small amount of left–lateral strike–slip, corresponding to a moment magnitude of Mw7.1, suggesting that the fault responsible for the earthquake was the northern segment of the DMCF (Deng Me Cuo Fault). The slip distribution results obtained from the finite fault model inversion show that this earthquake led to a significant increase in Coulomb stress at both ends of the fault and in the northeastern–southwestern region, with stress loading far exceeding the earthquake triggering threshold of 0.03 MPa. Through analysis, we believe that this Dingri earthquake occurred at the intersection of a “Y”-shaped structural feature where stress concentration is likely, which may be a primary reason for the frequent occurrence of moderate to strong earthquakes in this area. Full article
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<p>Tectonic background of the <span class="html-italic">M</span><sub>w</sub>7.1 Dingri earthquake in 2025. (<b>a</b>) Geographical Location of the Epicenter of the Dingri Earthquake. Black box: research sope of (<b>b</b>); gray focal spheres: source mechanisms of M &gt; 6.5 earthquakes since 1970 as provided by the USGS; Red focal sphere: source mechanism of the Dingri earthquake as given by the USGS. (<b>b</b>) Coverage Area of Earthquake Epicenter Images. Green and white boxes indicate the coverage areas of the European Space Agency’s Sentinel-1A ascending and descending orbits; black box: research sope of (<b>c</b>). (<b>c</b>) Local amplification map of the earthquake-prone area. The red and blue focal spheres represent the source mechanisms provided by the USGS and GCMT, respectively; the yellow dots: the precise aftershock catalog of the <span class="html-italic">M</span><sub>w</sub>7.1 Dingri earthquake [<a href="#B7-remotesensing-17-00936" class="html-bibr">7</a>]; white circles: the county towns in the vicinity of the epicenter; DMCF: Deng Me Cuo Fault; NHF: North Himalayan fault; YLZBF: Yarlung Zangbo River fault; SZDJF: Shenzha–Dingjie fault; TDF: Tangyako–Dingri fault.</p>
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<p>InSAR co-seismic deformation field of the <span class="html-italic">M</span><sub>w</sub>7.1 earthquake in January 2024: (<b>a</b>) T12 ascending track interferogram; (<b>b</b>) T12 ascending track displacement field; (<b>c</b>) T121 descending track interferogram; (<b>d</b>) T121 descending track displacement field; (<b>e</b>) T48 descending track interferogram; (<b>f</b>) T48 descending track displacement field. The red and blue focal spheres represent the source mechanism solutions of the January earthquake determined by the USGS and GCMT, respectively. The black fault traces indicate the identified active fault systems in the study area. One color fringe represents a line-of-sight (LOS) displacement of 50 mm. The red line segment indicates profile AA’, with the profile measurement results shown in <a href="#remotesensing-17-00936-f003" class="html-fig">Figure 3</a>. The red five-pointed star indicates the epicenter provided by the CEA. DMCF: Deng Me Cuo fault; NHF: North Himalayan fault; YLZBF: Yarlung Zangbo River fault; SZDJF: Shenzha–Dingjie fault; TDF: Tangyako–Dingri fault.</p>
Full article ">Figure 2 Cont.
<p>InSAR co-seismic deformation field of the <span class="html-italic">M</span><sub>w</sub>7.1 earthquake in January 2024: (<b>a</b>) T12 ascending track interferogram; (<b>b</b>) T12 ascending track displacement field; (<b>c</b>) T121 descending track interferogram; (<b>d</b>) T121 descending track displacement field; (<b>e</b>) T48 descending track interferogram; (<b>f</b>) T48 descending track displacement field. The red and blue focal spheres represent the source mechanism solutions of the January earthquake determined by the USGS and GCMT, respectively. The black fault traces indicate the identified active fault systems in the study area. One color fringe represents a line-of-sight (LOS) displacement of 50 mm. The red line segment indicates profile AA’, with the profile measurement results shown in <a href="#remotesensing-17-00936-f003" class="html-fig">Figure 3</a>. The red five-pointed star indicates the epicenter provided by the CEA. DMCF: Deng Me Cuo fault; NHF: North Himalayan fault; YLZBF: Yarlung Zangbo River fault; SZDJF: Shenzha–Dingjie fault; TDF: Tangyako–Dingri fault.</p>
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<p>Measurement results of co-seismic deformation profile of <span class="html-italic">M</span><sub>w</sub>7.1 Dingri earthquake in 2025. AA’ deformation field profile is shown in <a href="#remotesensing-17-00936-f002" class="html-fig">Figure 2</a>b,d.</p>
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<p>Estimation of fault geometric parameters. The red line in the histogram and the red dots in the 2D correlation plot indicate the maximum a posteriori solution.</p>
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<p>Inversion model of co-seismic slip distribution based on InSAR data. (<b>a</b>) Three-dimensional co-seismic slip distribution model; (<b>b</b>) surface projection of co-seismic slip distribution. Blue arrows indicate the slip direction.</p>
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<p>Fit of the inversion data for fault slip distribution. (<b>a</b>–<b>c</b>) correspond to the observed values, simulated values, and residuals of the ascending track InSAR data, respectively; (<b>d</b>–<b>f</b>) correspond to the observed values, simulated values, and residuals of the descending track InSAR data, respectively. The red rectangle is the fault plane projected on the surface; the black lines are active faults; DMCF: Deng Me Cuo fault; NHF: North Himalayan fault; SZDJF: Shenzha–Dingjie fault.</p>
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<p>Static ∆CFS in neighboring regions induced by the 2025 Dingri earthquake. (<b>a</b>) ∆CFS at a depth of 5 km underground. (<b>b</b>) ∆CFS at a depth of 7.5 km underground. (<b>c</b>) ∆CFS at a depth of 10 km underground. (<b>d</b>) ∆CFS at a depth of 5 km underground. The green dots the precise aftershock catalog of the <span class="html-italic">M</span><sub>w</sub>7.1 Dingri earthquake [<a href="#B7-remotesensing-17-00936" class="html-bibr">7</a>]; the black lines represent active faults. DMCF: Deng Me Cuo fault; NHF: North Himalayan fault; YLZBF: Yarlung Zangbo River fault; SZDJF: Shenzha–Dingjie Fault; TDF: Tangyako–Dingri Fault.</p>
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30 pages, 8502 KiB  
Article
Dynamic Structural Behavior of Monopile Support Structure for 15 MW Offshore Wind Turbine During Different Phases of Operation
by Sajid Ali, Muhammad Waleed and Daeyong Lee
J. Mar. Sci. Eng. 2025, 13(3), 515; https://doi.org/10.3390/jmse13030515 - 6 Mar 2025
Abstract
The structural integrity of offshore wind turbine monopiles is critical for ensuring operational stability and long-term performance under varying environmental and aerodynamic loads. However, transient load conditions during different operational phases, such as start, normal stop, and emergency stop, can significantly impact structural [...] Read more.
The structural integrity of offshore wind turbine monopiles is critical for ensuring operational stability and long-term performance under varying environmental and aerodynamic loads. However, transient load conditions during different operational phases, such as start, normal stop, and emergency stop, can significantly impact structural behavior, influencing fatigue life and dynamic stability. This study investigates the dynamic structural response of a 15 MW offshore wind turbine monopile, incorporating modal analysis and transient simulations to assess deflection, forces, moments, and rotational displacements at the mud-line. The modal analysis revealed natural frequencies of 0.509492 Hz, 1.51616 Hz, and 3.078425 Hz for the blade’s flap-wise modes, while side-to-side modes for the combined tower and monopile structure were identified at 0.17593 Hz, 0.922308 Hz, and 1.650862 Hz. These frequencies are crucial in evaluating resonance risks and ensuring dynamic stability under combined aerodynamic and hydrodynamic forces. The transient analysis demonstrated that lateral force (Fy) variations peaked at −2500 kN during emergency stop, while moment fluctuations (My) reached ±100,000 kNm, reflecting the monopile’s high dynamic sensitivity under sudden aerodynamic unloading. Rotational displacements also showed significant variations, with θx oscillating up to ±0.0009 degrees and θy between −0.0022 and −0.0027 degrees. These findings provide valuable insights into optimizing monopile design to mitigate resonance effects, improve fatigue performance, and enhance structural resilience for large-scale offshore wind turbine support systems. Full article
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<p>A 3D numerical model of NREL IEA 15 MW wind turbine in BLADED: (<b>a</b>) full model; (<b>b</b>) tower only (Note: blue layer is water and brown layer is mud-line in (<b>a</b>)).</p>
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<p>Estimation of the <span class="html-italic">EWS</span>: (<b>a</b>) Gumbel plot; (<b>b</b>) 50-year return gust.</p>
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<p>Weibull distribution for measured wind speed.</p>
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<p>Environmental loads with direction: (<b>a</b>) wind speed at hub height; (<b>b</b>) significant wave height.</p>
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<p>Monthly wave cycle at Buan (measured data).</p>
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<p>Modal analysis results for single blade.</p>
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<p>Modal analysis results for tower and monopile combined.</p>
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<p>Tower’s lateral oscillation amplitude for different phases of operation.</p>
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<p>Lateral (Fy) force on monopile at mud-line (red dot shows the exact location).</p>
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<p>Lateral (Fx) force on monopile at mud-line (red dot shows the exact location).</p>
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<p>Lateral (My) moment on monopile at mud-line (red dot shows the exact location).</p>
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<p>Lateral (Mx) moment on monopile at mud-line (red dot shows the exact location).</p>
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<p>Lateral (x-direction) deflection along the entire length of the tower and monopile with respect to time: (<b>a</b>) start; (<b>b</b>) normal stop; (<b>c</b>) emergency stop; (<b>d</b>) parked.</p>
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<p>Lateral (x-direction) deflection along the entire length of the tower and monopile with respect to time: (<b>a</b>) start; (<b>b</b>) normal stop; (<b>c</b>) emergency stop; (<b>d</b>) parked.</p>
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<p>Angular rotation (about <span class="html-italic">x</span>-axis) on monopile at mud-line (red dot shows the exact location).</p>
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<p>Angular rotation (about <span class="html-italic">y</span>-axis) on monopile at mud-line (red dot shows the exact location).</p>
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<p>Lateral force relative difference between parked and other states of operation: (<b>a</b>) Fx; (<b>b</b>) Fy.</p>
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<p>Moment relative difference between parked and other states of operation: (<b>a</b>) Mx; (<b>b</b>) My.</p>
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<p>Angular displacement relative difference between parked and other states of operation: (<b>a</b>) θx; (<b>b</b>) θy.</p>
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29 pages, 8917 KiB  
Article
Study on Eccentric Compression Behavior of Precast Stratified Concrete Composite Column with Inserted Steel Tube
by Yilin Wang, Shikun Ma and Shunyao Wang
Buildings 2025, 15(5), 826; https://doi.org/10.3390/buildings15050826 - 5 Mar 2025
Viewed by 99
Abstract
In order to improve the technical economy of steel-reinforced concrete structures and to promote the development of prefabricated concrete structures, a new type of partial precast steel-reinforced concrete column (precast stratified concrete composite column with inserted steel tube, PSCCST column) was proposed and [...] Read more.
In order to improve the technical economy of steel-reinforced concrete structures and to promote the development of prefabricated concrete structures, a new type of partial precast steel-reinforced concrete column (precast stratified concrete composite column with inserted steel tube, PSCCST column) was proposed and studied in this paper. Six PSCCST column specimens were tested to investigate their behavior under eccentric loading. The failure state, ultimate bearing capacities, load–strain relationship, as well as load-deflection curves were emphatically investigated. The failure modes of the PSCCST columns under eccentric compression and corresponding bearing capacity Nu calculation methods were proposed based on experimental research and analysis. The results of the study indicated that there are three main failure modes, which are compressive-type failure mode, total-yield-type failure mode, and tensile-type failure mode. The first two modes are preferable due to their more effective material utilization. The Nu of the PSCCST column was found to decrease obviously with the increase of eccentricity e. The deformation capacity denoted by the horizontal lateral deflection corresponding to Nu increased with the increase of e. Moreover, the proposed Nu calculation methods were proven to have high accuracy by the comparison with the experimental results (the average ratio of the calculated values to the experimental values was 0.95). Full article
(This article belongs to the Section Building Structures)
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<p>Schematic diagram of the PSCCST column.</p>
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<p>Section diagram of the PSCCST column under eccentric load.</p>
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<p>Schematic of the specimen.</p>
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<p>Layout of the loading device and measuring instruments.</p>
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<p>Sketch of the loading test frame and load application method.</p>
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<p>Final failure state of all specimens.</p>
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<p>Concrete strain distribution status.</p>
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<p>Steel tube strain distribution status.</p>
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<p>Lateral deflections at different load levels for all specimens.</p>
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<p>Vertical axial load versus lateral deflection of the mid-height section.</p>
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<p>Position of the neutral axis and the strain expression of each main part when <span class="html-italic">h</span><sub>2</sub> + <span class="html-italic">t</span><sub>1</sub> ≤ <span class="html-italic">x</span><sub>c</sub> &lt; <span class="html-italic">h</span><sub>0</sub>. <span class="html-italic">ε</span><sub>c</sub> is the concrete strain (under compression); <span class="html-italic">ε</span><sub>cu</sub> is the ultimate compressive strain of concrete, and <span class="html-italic">ε</span><sub>cu</sub> = 0.0033 is suitable for concrete, whose strength grade is not greater than C50, including the strength range of the concrete in the PSCCST columns; <span class="html-italic">ε</span>′<sub>su</sub> is the compression strain of the upper longitudinal reinforcement bars; <span class="html-italic">ε</span><sub>sl</sub> is the tensile strain of the lower reinforcement bars; <span class="html-italic">ε</span>′<sub>au</sub> is the compression strain of the upper flange; <span class="html-italic">ε</span>′a<sub>l</sub> is the compression strain of the lower flange.</p>
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<p>The neutral axis and strain expression of each main part when <span class="html-italic">h</span><sub>0</sub> ≤ <span class="html-italic">x</span><sub>c</sub> &lt; <span class="html-italic">h</span><sub>1</sub>. <span class="html-italic">ε</span>′<sub>sl</sub> is the compression strain of the lower longitudinal reinforcement bars.</p>
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<p>Position of the neutral axis and the strain expression of each main part when <span class="html-italic">x</span><sub>t</sub> ≤ <span class="html-italic">x</span><sub>c</sub> &lt; <span class="html-italic">h</span><sub>3</sub> + <span class="html-italic">t</span><sub>1</sub> + <span class="html-italic">t</span><sub>2</sub>. <span class="html-italic">ε</span><sub>al</sub> represents the tensile strain of the steel tube lower flange.</p>
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<p>The neutral axis and strain expression of each main part when <span class="html-italic">h</span><sub>3</sub> + <span class="html-italic">t</span><sub>1</sub> + <span class="html-italic">t</span><sub>2</sub> ≤ <span class="html-italic">x</span><sub>c</sub> &lt; <span class="html-italic">h</span><sub>2</sub> + <span class="html-italic">t</span><sub>1</sub>.</p>
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<p>Example of the strain and stress state of the normal section according to the plane section assumption.</p>
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<p>The stress and strain status of the normal section for Sub-case 1.</p>
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<p>The stress and strain statuses of the normal section (<span class="html-italic">h</span><sub>2</sub> + <span class="html-italic">t</span><sub>1</sub> ≤ <span class="html-italic">x</span><sub>c</sub> &lt; <span class="html-italic">h</span><sub>0</sub>).</p>
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<p>Strain and stress statuses of the normal section when <span class="html-italic">h</span><sub>3</sub> + <span class="html-italic">t</span><sub>1</sub> + <span class="html-italic">t</span><sub>2</sub> ≤ <span class="html-italic">x</span><sub>c</sub>&lt; <span class="html-italic">h</span><sub>2</sub> + <span class="html-italic">t</span><sub>1</sub>.</p>
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<p>Strain and stress of the normal section when <span class="html-italic">x</span><sub>t</sub> ≤ <span class="html-italic">x</span><sub>c</sub> &lt; <span class="html-italic">h</span><sub>3</sub> + <span class="html-italic">t</span><sub>1</sub> + <span class="html-italic">t</span><sub>2</sub>.</p>
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<p>Strain diagram of the limit state.</p>
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<p>The strain and stress of the normal section in the total-yield-type failure mode.</p>
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<p>Strain diagram of the limit state for the total-yield-type failure mode.</p>
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<p>Schematic diagram of the whole section of the steel tube under tension.</p>
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<p>Strain diagram of the limit state of the tensile-type failure mode.</p>
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<p>Strain and stress status of the tensile-type failure mode.</p>
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28 pages, 10856 KiB  
Article
Compressive Behavior of Long Simple and Multi-Cell CFT Columns When Using Tie Bars Connector Elements
by Nima Pahlavannejad Tabarestani, Morteza Naghipour and Stephen J. Hicks
Buildings 2025, 15(5), 817; https://doi.org/10.3390/buildings15050817 - 4 Mar 2025
Viewed by 286
Abstract
Concrete-filled steel tube (CFT) columns are increasingly used in high-rise structures due to their improved resilience to lateral loads. However, the behavior of multi-cell CFT columns, connected with different tie bar spacings, has been under-considered. This study aims to investigate the performance of [...] Read more.
Concrete-filled steel tube (CFT) columns are increasingly used in high-rise structures due to their improved resilience to lateral loads. However, the behavior of multi-cell CFT columns, connected with different tie bar spacings, has been under-considered. This study aims to investigate the performance of simple and four-cell CFT columns with tie bars at different spacings. Seven columns with different tie bar spacings (100, 300, and 500 mm) were examined under axial compression. The load–displacement curve, failure pattern, and concrete core failure characteristics were described. A calculation model for the axial pressure field of simple and four-cell CFT columns with tie bars at varying spacings was constructed using a finite element analysis software. The results showed that the axial compression load capacity of confined CFT columns was significantly higher in four-cell composite specimens, where the capacity increased by 15.6% and 33% with tie bar spacings of 500 mm and 300 mm, respectively. Also, compared to simple CFT specimens, the capacity increased by 14.7%, 27.8%, and 42.6% with tie bar spacings of 100 mm, 300 mm, and 500 mm, respectively. Full article
(This article belongs to the Special Issue Advances in Steel-Concrete Composite Structure—2nd Edition)
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<p>Simple and four-cell columns.</p>
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<p>Tensile test of the steel tube.</p>
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<p>Details of the test setup. (<b>a</b>) Columns and instrumentation, (<b>b</b>) location of the LVDTs and strain gauges.</p>
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<p>Failure modes of the simple CFT columns. (<b>a</b>) CFT column without tie bars, (<b>b</b>) CFT column with tie bars.</p>
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<p>Failure modes of the columns with tie bars with a 300 mm spacing.</p>
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<p>Failure modes (<b>a</b>) and force–displacement curve (<b>b</b>) of simple CFT columns with different tie bars.</p>
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<p>Failure modes of the four-cell CFT columns with tie bars with a 500 mm spacing.</p>
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<p>Failure modes (<b>a</b>) and force–displacement curve (<b>b</b>) of four-cell CFT columns with tie bars in different spacing.</p>
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<p>Force–displacement curve of simple and four-cell CFT columns with tie bars with different spacings.</p>
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<p>Force–strain curve of simple and four-cell CFT columns.</p>
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<p>Mesh types used in numerical modeling. (<b>a</b>) Simple CFT model without tie bars, (<b>b</b>) simple CFT model with tie bars with a 100 mm spacing, (<b>c</b>) four-cell CFT model with tie bars with a 100 mm spacing.</p>
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<p>Boundary and loading conditions.</p>
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<p>Stress–strain curves of the steel column for steel thicknesses of (<b>a</b>) 2 mm and (<b>b</b>) 4 mm.</p>
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<p>Stress–strain curves of confined and unconfined concrete.</p>
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<p>Diagram for confining pressure (with tie bars).</p>
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<p>Comparison of the numerical and experimental results.</p>
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<p>Comparison of failure in experimental and numerical models. (<b>a</b>) Simple CFT column without tie bars, (<b>b</b>) simple CFT column with 500 mm spaced tie bars, (<b>c</b>) simple CFT column with 500 mm spaced tie bars.</p>
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<p>Load–displacement curves of simple and four-cell CFT columns with a 2 mm steel tube and tie bars. (<b>a</b>) Compressive strength of 20 MPa, (<b>b</b>) compressive strength of 40 MPa, and (<b>c</b>) compressive strength of 80 MPa.</p>
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<p>Load–displacement curves of simple and four-cell CFT columns with 4 mm steel tube and tie bars. (<b>a</b>) Compressive strength of 20 MPa, (<b>b</b>) compressive strength of 40 MPa, and (<b>c</b>) compressive strength of 80 MPa.</p>
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<p>Concrete contribution ratio for the experimental test with different spacings of tie bars.</p>
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<p>Concrete contribution ratio for the numerical test under different D/t ratios, (<b>a</b>) simple CFT columns, (<b>b</b>) four-cell CFT Columns.</p>
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<p>Strength index (SI) for the experimental test with different spacings of the tie bars.</p>
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<p>Strength index (SI) for the numerical test under different D/t ratios. (<b>a</b>) Simple CFT columns, (<b>b</b>) four-cell CFT columns.</p>
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<p>Concrete confinement of the CFT columns with different D/t ratios. (<b>a</b>) Simple CFT columns, (<b>b</b>) four-cell CFT columns.</p>
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<p>Ductility index of the CFT columns with different D/t ratios. (<b>a</b>) Simple CFT columns, (<b>b</b>) four-cell CFT columns.</p>
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24 pages, 7568 KiB  
Article
Delayed Detached-Eddy Simulations of Aerodynamic Variability During Carrier-Based Aircraft Landing with a Domain Precursor Inflow Method
by Jiawei Fu, Ruifan Hu, Hong Wang, Ke Xu and Shuling Tian
J. Mar. Sci. Eng. 2025, 13(3), 498; https://doi.org/10.3390/jmse13030498 - 3 Mar 2025
Viewed by 151
Abstract
Flight tests and wind tunnel experiments face difficulties in investigating the impact of aircraft carrier air-wake on the landing process. Meanwhile, numerical methods generally exhibit low overall computational efficiency in solving such problems. To address the computational challenges posed by the disparate spatiotemporal [...] Read more.
Flight tests and wind tunnel experiments face difficulties in investigating the impact of aircraft carrier air-wake on the landing process. Meanwhile, numerical methods generally exhibit low overall computational efficiency in solving such problems. To address the computational challenges posed by the disparate spatiotemporal scales of the ship air-wake and aircraft motion, a domain precursor inflow method is developed to efficiently generate unsteady inflow boundary conditions from precomputed full-domain air-wake simulations. This study investigates the aerodynamic variability of carrier-based aircraft during landing through the turbulent air-wake generated by an aircraft carrier, employing a hybrid RANS-LES methodology on dynamic unstructured overset grids. The numerical framework integrates a delayed detached-eddy simulation (DDES) model with a parallel dynamic overset grid approach, enabling high-fidelity simulations of coupled aircraft carrier interactions. Validation confirms the accuracy of the precursor inflow method in reproducing air-wake characteristics and aerodynamic loads compared to full-domain simulations. Parametric analyses of 15 distinct landing trajectories reveal significant aerodynamic variability, particularly within 250 m of the carrier, where interactions with island-generated vortices induce fluctuations in lift (up to 25%), drag (18%), and pitching moments (30%). Ground effects near the deck further amplify load variations, while lateral deviations in landing paths generate asymmetric forces and moments. The proposed methodology demonstrates computational efficiency for multi-scenario analysis, providing critical insights into aerodynamic uncertainties during carrier operations. Full article
(This article belongs to the Section Ocean Engineering)
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<p>The hole cutting of the current overset grid approach. (<b>a</b>) Definition of the active zone for each body; (<b>b</b>) classification of grid nodes based on their wall distance and (<b>c</b>) hole cutting result for 4 circles.</p>
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<p>Definition of interpolation nodes and parallel donor cell search for interpolation nodes. (<b>a</b>) Double-layer interpolation boundary in the overset grid; (<b>b</b>) determination of candidate donor grid parts in the candidate grid with oriented bounding boxes and (<b>c</b>) searching the donor cells in the candidate donor parts in their processor.</p>
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<p>Schematic of carrier-based aircraft landing. (<b>a</b>) Flight path; (<b>b</b>) velocity components.</p>
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<p>Schematic of the domain precursor inflow simulation approach. (<b>a</b>) The entire domain of ship air-wake simulation with a velocity inflow boundary condition; (<b>b</b>) the aft air-wake domain with a precursor inflow boundary condition.</p>
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<p>The configurations of the carrier model and carrier-based aircraft. (<b>a</b>) Front view of the aircraft carrier model CVN; (<b>b</b>) lateral view of the aircraft carrier model CVN; (<b>c</b>) CVN with the aft portion of the carrier deck selected for landing simulation; (<b>d</b>) front view of a simplified carrier-based aircraft model; (<b>e</b>) lateral view of the aircraft model; (<b>f</b>) bottom view of the aircraft model.</p>
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<p>Grids for air-wake field computations and aircraft landing simulations. (<b>a</b>) Global distribution and (<b>b</b>) local detail of the grid in the entire air-wake field; (<b>c</b>) grid in the aft air-wake field and (<b>d</b>) grid of the aircraft.</p>
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<p>The vortex structures illustrated with the iso-surface of the Q-criterion: (<b>a</b>) result of RANS; (<b>b</b>) result of the DDES-based LES-RANS hybrid method.</p>
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<p>The positions of four probe points.</p>
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<p>The velocity fluctuations calculated by using the two methods at (<b>a</b>) probe point 1, (<b>b</b>) probe point 2, (<b>c</b>) probe point 3, and (<b>d</b>) probe point 4.</p>
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<p>The instantaneous vortex structures realized by displaying the iso-surface of the Q-criterion at (<b>a</b>) 50 s, (<b>b</b>) 60 s, and (<b>c</b>) 70 s calculated by using (<b>i</b>) the entire domain simulation method and (<b>ii</b>) the domain precursor inflow method.</p>
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<p>Comparison of (<b>a</b>) force coefficient and (<b>b</b>) moment coefficient during the landing process of a carrier-based aircraft.</p>
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<p>The instantaneous iso-surface of the Q-criterion at (<b>a</b>) 44 s, (<b>b</b>) 48 s, and (<b>c</b>) 52 s calculated by using (<b>i</b>) the entire domain simulation method and (<b>ii</b>) the domain precursor inflow method.</p>
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<p>The overset grid of the carrier-based aircraft at (<b>a</b>) 200 m and (<b>b</b>) 50 m.</p>
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<p>Landing paths utilized in the computations to demonstrate the aerodynamic variability.</p>
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<p>The positions of the aircraft’s center of mass during landings. (<b>a</b>) Longitudinal coordination varying with time, (<b>b</b>) vertical coordination varying with time, and (<b>c</b>) lateral coordination varying with time.</p>
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<p>Force and moment coefficients during landing. (<b>a</b>) Lift coefficient, (<b>b</b>) drag coefficient, (<b>c</b>) lateral force coefficient, (<b>d</b>) rolling moment coefficient, (<b>e</b>) pitching moment coefficient, and (<b>f</b>) yawing moment coefficient.</p>
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<p>The uncertainty of force and moment coefficients during landing. (<b>a</b>) Lift coefficient, (<b>b</b>) drag coefficient, (<b>c</b>) lateral force coefficient, (<b>d</b>) rolling moment coefficient, (<b>e</b>) pitching moment coefficient, and (<b>f</b>) yawing moment coefficient.</p>
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<p>The air-wake flow field. (<b>a</b>) Vortex structures and (<b>b</b>) velocity distribution and streamlines in the vertical plane where the landing glide slope is located.</p>
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<p>The coupled flow field of the carrier’s air-wake and the aircraft field when the distance between the carrier-based aircraft and the landing point is (<b>a</b>) 475 m, (<b>b</b>) 375 m, (<b>c</b>) 200 m, (<b>d</b>) 125 m, (<b>e</b>) 50 m, and (<b>f</b>) 0 m.</p>
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<p>The distributions of the pressure and streamline in the vertical plane for the carrier-based aircraft at the longitudinal positions of (<b>a</b>) 200 m, (<b>b</b>) 125 m, (<b>c</b>) 50 m, and (<b>d</b>) 0 m.</p>
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<p>The instantaneous vortex structure illustrated with the iso-surface of the Q-criterion around the carrier-based aircraft at the longitudinal positions of (<b>a</b>) 200 m, (<b>b</b>) 125 m, (<b>c</b>) 50 m, and (<b>d</b>) 0 m.</p>
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35 pages, 760 KiB  
Article
A Comparison of Three Theories for Vibration Analysis for Shell Models
by Maria Anna De Rosa, Isaac Elishakoff and Maria Lippiello
CivilEng 2025, 6(1), 13; https://doi.org/10.3390/civileng6010013 - 3 Mar 2025
Viewed by 132
Abstract
Shells are significant structural components that are extensively utilized in numerous engineering fields, including architectural and infrastructural projects. These components are employed in the construction of domes, water tanks, stadiums and auditoriums, hangars, and cooling towers. Significant research efforts have been dedicated to [...] Read more.
Shells are significant structural components that are extensively utilized in numerous engineering fields, including architectural and infrastructural projects. These components are employed in the construction of domes, water tanks, stadiums and auditoriums, hangars, and cooling towers. Significant research efforts have been dedicated to the analysis of vibrations and dynamic behaviors of shells, due to their distinctive capacity to efficiently bear loads through their geometry rather than mass. Additionally, a vast array of shell theories and computational methods have been proposed and developed by researchers. This paper represents a continuation of research initiated begun in a 2009 paper by Elishakoff, wherein the suggestion was made to disregard an energetic term in the dynamic analysis of Timoshenko–Ehrenfest beams, wherein the suggestion was made to disregard an energetic term in the dynamic analysis of Timoshenko–Ehrenfest beams. The resulting reduced theory was found to be both more straightforward and more reliable than the complete, classical approach. While the original idea was heuristically justified, a more sound variationally consistent theory was proposed in the papers of De Rosa et al. concerning the dynamic analysis of the Timoshenko-Ehrenfest beams and later extended to the case of the Uflyand-Mindlin plates. In accordance with the proposal put forth in those works, we initially delineate the classical shell theory and subsequently propose two alternative hypotheses that give rise to two distinct aspects of the energy terms. By employing the variational approach, we derive two novel boundary problems, which are direct generalizations of those previously considered. Both theories can be readily specialized for beams and plates, and the theory can also be specialized for the case of cylindrical shells. Full article
(This article belongs to the Section Mathematical Models for Civil Engineering)
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<p>Reference shell element [<a href="#B5-civileng-06-00013" class="html-bibr">5</a>].</p>
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<p>Boundary force and moment resultants [<a href="#B5-civileng-06-00013" class="html-bibr">5</a>].</p>
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<p>Simply supported rectangular beam.</p>
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<p>Rectangular plate with four simply supported edges.</p>
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<p>Coordinate definitions for a circular cylindrical shell [<a href="#B5-civileng-06-00013" class="html-bibr">5</a>].</p>
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15 pages, 937 KiB  
Article
Incorporating Non-Linear Epoxy Resin Development in Infusion Simulations: A Dual-Exponent Viscosity Model Approach
by Mohammad W. Tahir, Umar Khan and Jan-Peter Schümann
Polymers 2025, 17(5), 657; https://doi.org/10.3390/polym17050657 - 28 Feb 2025
Viewed by 168
Abstract
In the field of liquid composite moulding (LCM) simulations, a long-standing assumption has dominated–the belief in constant resin viscosity. While effective in many cases, this assumption may not hold for the infusion process, which lasts for an extended period. This impacts the mechanical [...] Read more.
In the field of liquid composite moulding (LCM) simulations, a long-standing assumption has dominated–the belief in constant resin viscosity. While effective in many cases, this assumption may not hold for the infusion process, which lasts for an extended period. This impacts the mechanical properties of the cured epoxy, which are crucial for load transfer in polymer structures. The majority of epoxy resins operate on a bipartite foundation, wherein their viscosity undergoes dynamic alterations during the process of cross-linking. Temperature and cross-linking intricately interact, with elevated temperatures initially reducing viscosity due to kinetic energy but later increasing it as cross-linking accelerates. This interplay significantly influences the efficiency of the infusion process, especially in large and intricate moulds. This article explores the significant temperature dependence of epoxy resin viscosity, proposing an accurate model rooted in its non-linear evolution. This model aligns with empirical evidence, offering insights into determining the optimal starting temperature for efficient mould filling. This study presents an advanced infusion model that extends existing non-linear dual-split viscosity approaches by incorporating the experimental validation of viscosity variations. Unlike previous models that primarily focus on theoretical or numerical frameworks, this work integrates experimental insights to optimize infusion temperature for efficient resin infusion in large and complex parts. Building on these findings, a novel mould-filling technique is proposed to enhance efficiency and reduce material waste. Full article
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<p>Schematic of the LCM.</p>
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<p>Plot of temperature development of RIMR 035c epoxy resin with four different hardeners redrawn from manufacturer datasheet.</p>
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<p>Plot of viscosity development with curing for the epoxy resin with fitted data for single exponent model presented in Equation (<a href="#FD1-polymers-17-00657" class="html-disp-formula">1</a>).</p>
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<p>Plot of viscosity development with curing for the epoxy resin with fitted data for dual exponent model shown in Equation (<a href="#FD3-polymers-17-00657" class="html-disp-formula">3</a>).</p>
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<p>Schematic diagram of a typical infusion process in Equation (<a href="#FD3-polymers-17-00657" class="html-disp-formula">3</a>).</p>
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<p>Plot of temperature development of RIMR 035c epoxy resin with four different hardeners under insulated setup [<a href="#B29-polymers-17-00657" class="html-bibr">29</a>].</p>
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<p>Comparison of the flow front position for real infusion of epoxy and the model presented in Equation (<a href="#FD6-polymers-17-00657" class="html-disp-formula">6</a>).</p>
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<p>Plot of the flow front with time for UD flow case. The dotted red line exemplifies a three-hour mould filling time-frame. Infusion at 40 or 45 °C results in an early gelation point causing an incomplete mould filling, while 25 °C causes inefficient filling and extended duration. An optimal starting temperature of 35 °C ensures efficient mould filling.</p>
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27 pages, 23631 KiB  
Article
Traditional Malay House Preservation: Guidelines for Structural Evaluation
by Sara Alsheikh Mahmoud and Huzaifa Bin Hashim
Buildings 2025, 15(5), 782; https://doi.org/10.3390/buildings15050782 - 27 Feb 2025
Viewed by 282
Abstract
The traditional Malay house is a significant component of the Malay cultural heritage and a key example of vernacular architecture. It is characterised by its outstanding design and the various styles across Malaysia. Traditional Malay houses experience deterioration and damage due to various [...] Read more.
The traditional Malay house is a significant component of the Malay cultural heritage and a key example of vernacular architecture. It is characterised by its outstanding design and the various styles across Malaysia. Traditional Malay houses experience deterioration and damage due to various threats, resulting in many houses being abandoned. A thorough structural evaluation is crucial for preserving the traditional Malay house. This research aimed to develop guidelines for the global structural evaluation of the Malay house. A case study approach was adopted in this research. Site visits, visual surveys, geometrical surveys, and dilapidation surveys were also employed. The research involved structural analysis using SAP2000. The results revealed the vulnerability of the houses to lateral forces, sliding, and differential settlement under scouring. The key structural members have adequate load-bearing capacity, which might be compromised under certain conditions, as in the case of deterioration. These results helped identify potential safety concerns and led to the development of guidelines for the global structural evaluation of Malay houses. The guidelines cover analysis inputs and modelling techniques in terms of material, geometry, joints, and foundations. They address load criteria and the impacts of flooding and scouring on the structural behaviour of the traditional Malay house. The guidelines, finally, recommend that structural checks be considered. This research contributes to traditional Malay house preservation by providing an evidence-based approach to designing preservation measures. Full article
(This article belongs to the Section Building Structures)
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<p>An example of a traditional Malay house.</p>
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<p>An example of tenon–mortise joints in TMH.</p>
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<p>Typical column–base connection in traditional Malay houses [<a href="#B35-buildings-15-00782" class="html-bibr">35</a>].</p>
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<p>Research methodology flowchart.</p>
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<p>Wan Muda Wan Hassan’s house.</p>
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<p>The three-dimensional model of the traditional Malay house case study.</p>
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<p>Joint displacement of the case study with respect to different material models.</p>
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<p>Joint displacement of the case study with respect to diminishing material stiffness.</p>
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<p>Changes in axial forces of a key column experiencing a reduction in the cross-section.</p>
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<p>Changes in stresses of a key column experiencing a reduction in the cross-section.</p>
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<p>Joint displacements of the case study with respect to diminishing joint stiffness.</p>
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<p>Settlement of the foundation under the combined effect of gravity, wind, and flood loads.</p>
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<p>Settlement of the foundation under the combined effect of loads (gravity, wind, flood) and scouring.</p>
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<p>Load path under gravity loads.</p>
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<p>Load path under the combined effects of gravity, flood, and wind loads.</p>
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<p>Sliding and resisting forces of the TMH under different loads and conditions.</p>
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<p>Key beams’ capacity vs. developed stresses under different loads and conditions.</p>
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<p>Key columns’ capacity vs. developed stresses under different loads and conditions.</p>
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13 pages, 2007 KiB  
Article
Comparison of Outcomes Between Functionally and Mechanically Aligned Total Knee Arthroplasty: Analysis of Parallelism to the Ground and Weight-Bearing Position of the Knee Using Hip-to-Calcaneus Radiographs
by Hongyeol Yang, Chanjin Park, Jaehyeok Cheon, Jaeyeon Hwang and Jongkeun Seon
J. Pers. Med. 2025, 15(3), 91; https://doi.org/10.3390/jpm15030091 - 27 Feb 2025
Viewed by 143
Abstract
Background: The objective of this study was to compare the outcomes between patients undergoing mechanically aligned conventional total knee arthroplasty (MA-CTKA) and functionally aligned robotic-arm-assisted TKA (FA-RTKA). Methods: We reviewed a prospectively collected database of consecutive patients who underwent primary total knee arthroplasty [...] Read more.
Background: The objective of this study was to compare the outcomes between patients undergoing mechanically aligned conventional total knee arthroplasty (MA-CTKA) and functionally aligned robotic-arm-assisted TKA (FA-RTKA). Methods: We reviewed a prospectively collected database of consecutive patients who underwent primary total knee arthroplasty (TKA) for knee osteoarthritis between June 2022 and May 2023. Patients were divided into two groups—MA-CTKA (n = 50) and FA-RTKA (n = 50)—based on the introduction of a robotic-arm-assisted system during the study period. The hip–knee–ankle (HKA) angle, joint line orientation angle (JLOA) relative to the floor, and weight-bearing line (WBL) ratio were evaluated using full-length standing hip-to-calcaneus radiographs to compare the conventional mechanical axis (MA) and the ground mechanical axis (GA) passing through the knee joint between the groups. Clinical outcomes were also compared between the two groups. Results: There were no significant differences in the postoperative HKA angle between the groups, due to discrepancies in the targeted alignment strategies (FA-RTKA: 2.0° vs. MA-CTKA: 0.5°; p = 0.001). The postoperative JLOA in the FA-RTKA group was more parallel to the floor, whereas the MA-CTKA group showed a downward angulation toward the lateral side (0.6° vs. −2.7°; p < 0.001). In the FA-RTKA group, the GA passed through a neutral position when accounting for the calcaneus, while the MA-CTKA group showed a more lateral GA position (48.8% vs. 53.8%; p = 0.001). No significant differences in clinical outcomes were shown between the FA-RTKA and MA-CTKA groups, with the FA-RTKA group demonstrating higher Forgotten Joint Scores and a greater range of motion (all p < 0.05). Conclusions: Functionally aligned TKA demonstrated improved joint line parallelism to the floor and more neutral weight-bearing alignment in the GA compared to mechanically aligned TKA. These findings indicate a more balanced load distribution across the knee, which may contribute to the superior clinical outcomes observed in the functionally aligned group. Full article
(This article belongs to the Section Methodology, Drug and Device Discovery)
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<p>Intraoperative component position using the functional alignment principle. The tibial component alignment was first set to 3.0° (varus) in a coronal alignment to maintain the native tibial joint line within the boundaries; subsequently, the extension and flexion gaps were balanced by fine-tuning the femoral component alignment in all three dimensions. The femoral component coronal alignment was set to −1.1° (valgus) and 3.9° externally rotated relative to the posterior condylar axis. The symbols represent the same reference point.</p>
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<p>Assessment of the JLOA on hip-to-calcaneus radiographs. (<b>A</b>) In the functionally aligned group, the postoperative JLOA was parallel to the ground. (<b>B</b>) In the mechanically aligned group, the postoperative JLOA was inclined downward toward the lateral side. The dotted line indicates the ground orientation (G). The JLOA was defined as the angle between the proximal tibial joint surface (solid red line) and the ground (dashed red line).</p>
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<p>Assessment of weight-bearing position in the knee joint on hip-to-calcaneus radiographs in functionally aligned total knee arthroplasty. (<b>A</b>) The Ground Mechanical Axis, defined as the line (GA; solid red) extending from the center of the femoral head to the lowest point of the calcaneus, runs lateral to the traditional mechanical axis (MA; dashed red line). (<b>B</b>) Under the “true” condition, when accounting for the calcaneus, the GA passes through the center of the knee joint. (<b>C</b>) In contrast, the traditional MA passes slightly medially through the knee joint. The ‘R’ symbol represents ‘Right’ in the X-ray marking.</p>
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35 pages, 12064 KiB  
Article
An Adaptive GPR-Based Multidisciplinary Design Optimization of Structural and Control Parameters of Intelligent Bus for Rollover Stability
by Tingting Wang, Xu Shao, Dongchen Qin, Kun Huang, Mingkuan Yao and Yuechen Duan
Mathematics 2025, 13(5), 782; https://doi.org/10.3390/math13050782 - 26 Feb 2025
Viewed by 280
Abstract
Considering the influence of high-speed obstacle avoidance trajectory in the optimization design stage of intelligent bus aerodynamic shape. A collaborative optimization method aiming at aerodynamic structure and trajectory control system for intelligent bus rollover stability is proposed to reduce the interference of lateral [...] Read more.
Considering the influence of high-speed obstacle avoidance trajectory in the optimization design stage of intelligent bus aerodynamic shape. A collaborative optimization method aiming at aerodynamic structure and trajectory control system for intelligent bus rollover stability is proposed to reduce the interference of lateral aerodynamic load caused by large bus side area on driving stability and improve the rollover safety of intelligent bus in high-speed obstacle avoidance process. At the conceptual design stage, a multidisciplinary co-design optimization frame of aerodynamics/dynamics/control is built, and an adaptive Gaussian Process Regression approximate modeling method is proposed to establish an approximate model of high-precision and high-efficiency rollover evaluation index with rollover stability as the optimization objective and obstacle avoidance safety and resistance to crosswind interference as constraints. Taking rollover stability and obstacle avoidance safety as the optimization objectives, the integrated design of static structural parameters and dynamic control parameters of intelligent buses is carried out. The results show that the proposed MDO method can obtain the aerodynamic shape of the vehicle body with low crosswind sensitivity and a safe and stable obstacle avoidance trajectory. Compared with the initial trajectory, the peak lateral load transfer rate during the obstacle avoidance process decreases by 33.91%, which significantly reduces the risk of rollover. Compared with the traditional serial optimization method, the proposed co-design optimization method has obvious advantages and can further improve the driving safety performance of intelligent buses. Full article
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<p>System level and subsystem level of MDO.</p>
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<p>Dynamic design variable discretization.</p>
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<p>Adaptive Gaussian modeling method process based on gene sequence.</p>
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<p>Adaptive Gaussian modeling method principle based on gene sequence.</p>
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<p>Aerodynamic/dynamic coupling model.</p>
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<p>Aerodynamic force (Numbers 1–6 are front top arc, front side arc, rear top arc, rear side arc, side arc, and tail upwarp).</p>
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<p>Steering wheel control method.</p>
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<p>Obstacle avoidance trajectory cluster.</p>
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<p>Relationship between yaw and LTR (The gray box indicates that in the process of high-speed obstacle avoidance, under this body attitude, both driving states are in critical rollover).</p>
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<p>CFD turbulence model selection.</p>
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<p>Cross-wind simulation method (The arrow shows the wind direction).</p>
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<p>Mesh generation.</p>
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<p>Encryption domain.</p>
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<p>Bus CFD results. (<b>a</b>) Bus surface pressure; (<b>b</b>) pressure distribution at 1200 mm above ground of a bus; (<b>c</b>) velocity distribution at 1200 mm above ground of the bus.</p>
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<p>Bus surface pressure at yaw angle of 14°.</p>
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<p>Intelligent bus MDO strategy.</p>
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<p>(<b>a</b>) LTR—Gaussian; (<b>b</b>) LTR—kriging; (<b>c</b>) LTR—neural network. blue areas represent a 95% confidence interval.</p>
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<p>LTR iterative curve.</p>
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<p>Control variables before and after optimization.</p>
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<p>Trajectories before and after optimization. Dashed lines indicate highway lane lines.</p>
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<p>Aerodynamic load coefficient optimization results.</p>
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<p>Optimization results.</p>
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<p>Aerodynamic force coefficient.</p>
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<p>Approximate result of adaptive combinatorial kernel function.</p>
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<p>Test function iteration comparison diagram. (<b>a</b>) F1 function; (<b>b</b>) F2 function; (<b>c</b>) F3 function.</p>
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15 pages, 2153 KiB  
Article
Horizontal Bearing Characteristics of Large-Diameter Rock-Socketed Rigid Pile and Flexible Pile
by Lin Liu, Li Xiao, Yang Liu, Mingrui Zhao, Fan Jin, Xiangyu Li and Yu Tian
Buildings 2025, 15(5), 768; https://doi.org/10.3390/buildings15050768 - 26 Feb 2025
Viewed by 191
Abstract
In order to study the horizontal bearing characteristics of large-diameter rock-socketed rigid pile and flexible pile, two lateral loading tests in which the pile lengths are 5.2 m and 11.07 m were carried out. Unidirectional multi-cyclic loading was applied to the piles during [...] Read more.
In order to study the horizontal bearing characteristics of large-diameter rock-socketed rigid pile and flexible pile, two lateral loading tests in which the pile lengths are 5.2 m and 11.07 m were carried out. Unidirectional multi-cyclic loading was applied to the piles during the tests, with the maximum load reaching 3500 kN. The measured results are compared with the calculated results of Zhang’s method, m-method and the rigid pile method in the design codes. It is indicated that if the characteristic values of the horizontal bearing capacity of the large-diameter rock-socketed rigid pile and flexible pile are determined by the same horizontal displacement of the pile head, some risk will be brought to the design of the rigid pile. Compared with the rigid pile method, the m-method is more suitable for calculating the rotation angle of the pile head. In terms of the maximum bending moment of the large-diameter rock-socketed flexible pile under the critical load, the calculated result of Zhang’s method is less than the measured result, while the calculated result of the m-method is the largest. However, for the rigid pile, both Zhang’s method and m-method underestimate the maximum bending moment of the pile body. In summary, when a large-diameter rock-socketed pile is designed, reasonable calculation method and failure discrimination standard should be chosen according to the actual conditions. Full article
(This article belongs to the Special Issue Dynamic Response of Civil Engineering Structures under Seismic Loads)
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<p>Typical profile of the test site.</p>
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<p>Diagram of the loading device. (<b>a</b>) Top view. (<b>b</b>) Front view.</p>
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<p>Sensors layout on the test pile.</p>
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<p>Loading curves of Pile 11#. (<b>a</b>) <span class="html-italic">H</span>–<span class="html-italic">t</span>–<span class="html-italic">y</span><sub>0</sub> curve. (<b>b</b>) <span class="html-italic">H</span>–∆<span class="html-italic">y</span><sub>0</sub>/∆<span class="html-italic">H</span> curve. (<b>c</b>) <span class="html-italic">H</span>–<span class="html-italic">σ</span><sub>s</sub> curve.</p>
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<p>Loading curves of Pile 13#. (<b>a</b>) <span class="html-italic">H</span>–<span class="html-italic">t</span>–<span class="html-italic">y</span><sub>0</sub> curve. (<b>b</b>) <span class="html-italic">H</span>–∆<span class="html-italic">y</span><sub>0</sub>/∆<span class="html-italic">H</span> curve. (<b>c</b>) <span class="html-italic">H</span>–<span class="html-italic">σ</span><sub>s</sub> curve.</p>
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<p>Rotation angle of the pile head. (<b>a</b>) Pile 11#. (<b>b</b>) Pile 13#.</p>
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<p>Measured bending moment of Pile 11# and Pile 13#. (<b>a</b>) 350 kN~1750 kN. (<b>b</b>) 2100 kN~3500 kN.</p>
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<p>Calculated bending moment of Pile 11#. (<b>a</b>) 350 kN~1750 kN. (<b>b</b>) 2100 kN~3500 kN.</p>
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<p>Calculated bending moment of Pile 13#. (<b>a</b>) 350 kN~1750 kN. (<b>b</b>) 2100 kN~3500 kN.</p>
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13 pages, 3156 KiB  
Article
Alterations in the Neuromuscular Control Mechanism of the Legs During a Post-Fatigue Landing Make the Lower Limbs More Susceptible to Injury
by Penglei Fan, Youngsuk Kim, Dong-Wook Han, Sukwon Kim and Ting Wang
Bioengineering 2025, 12(3), 233; https://doi.org/10.3390/bioengineering12030233 - 24 Feb 2025
Viewed by 293
Abstract
Fatigue causes the lower limb to land in an injury-prone state, but the underlying neuromuscular control changes remain unclear. This study aims to investigate lower limb muscle synergies during landing in basketball players, both before and after fatigue, to examine alterations in neuromuscular [...] Read more.
Fatigue causes the lower limb to land in an injury-prone state, but the underlying neuromuscular control changes remain unclear. This study aims to investigate lower limb muscle synergies during landing in basketball players, both before and after fatigue, to examine alterations in neuromuscular control strategies induced by fatigue. Eighteen male recreational basketball players performed landing tasks pre- and post-fatigue induced by 10 × 10 countermovement jumps. Electromyographic (EMG) data from eight muscles, including the erector spinae (ES), rectus abdominus (RA), gluteus maximus (GM), rectus femoris (RF), biceps femoris (BF), lateral gastrocnemius (LG), soleus (SM), and tibialis anterior (TA) muscles, were analyzed using non-negative matrix factorization to extract muscle synergies. Post-fatigue results revealed significant changes: synergy primitive 1 decreased before landing (18–30% phase) and synergy primitive 2 decreased after landing (60–100% phase). Muscle weights of the LG and SM in synergy module 2 increased. Fatigue reduced synergistic muscle activation levels, compromising joint stability and increasing knee joint loading due to greater reliance on calf muscles. These changes heighten the risk of lower limb injuries. To mitigate fatigue-induced injury risks, athletes should improve thigh muscle endurance and enhance neuromuscular control, fostering better synergy between thigh and calf muscles during fatigued conditions. Full article
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<p>Schematic diagram of the experimental flow.</p>
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<p>EMG sensor placement position and experimental setup. ES: erector spinae; RA: rectus abdominus; GM: gluteus maximus; RF: rectus femoris; BF: biceps femoris; LG: lateral gastrocnemius; SM: soleus; TA: tibialis anterior.</p>
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<p>Flowchart of EMG signal processing to obtain the EMG matrix. ES: erector spinae; RA: rectus abdominus; GM: gluteus maximus; RF: rectus femoris; BF: biceps femoris; LG: lateral gastrocnemius; SM: soleus; TA: tibialis anterior.</p>
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<p>Synergy modules and synergy primitives pre- and post-fatigue intervention in 18 participants (M1 is synergy module 1, M2 is synergy module 2, P1 is synergy primitive 1, and P2 is synergy primitive 2; mean ± SD). In the M-plot, the Y-axis is the weight of the muscle, and the X-axis is the muscle name. In the P-plot, the Y-axis represents the degree of activation of the module, and the X-axis represents the normalized 101 data points, with the 50% dashed line in the middle indicating the moment of initial contact with the ground.</p>
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<p>Paired <span class="html-italic">t</span>-test results for each muscle in synergy module 1 and synergy module 2 for pre and post-fatigue intervention. Pre-fatigue intervention in blue and post-fatigue intervention in red, * is a significant difference (<span class="html-italic">p</span> &lt; 0.05), ** is a very significant difference (<span class="html-italic">p</span> &lt; 0.01).</p>
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<p>SPM paired samples <span class="html-italic">t</span>-test comparing differences between synergy primitives in pre- and post-fatigue interventions. The blue line as pre-fatigue intervention, the red line as post-fatigue intervention, with shading denoting the stage in which the difference occurred.</p>
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22 pages, 6784 KiB  
Article
Effect of the Initial Damage State on the Seismic Behavior of A Five-Story CLT Building
by Franco Benedetti, Julieta Álvarez M., Alan Jara-Cisterna, Alexander Opazo-Vega and Víctor Rosales
Buildings 2025, 15(5), 727; https://doi.org/10.3390/buildings15050727 - 24 Feb 2025
Viewed by 168
Abstract
Timber construction experiences a growing trend in different countries due to its inherent environmental benefits and proven lateral load performance. However, most of the previous studies on structural and seismic performance have focused on undamaged structures without any signs of deterioration. This paper [...] Read more.
Timber construction experiences a growing trend in different countries due to its inherent environmental benefits and proven lateral load performance. However, most of the previous studies on structural and seismic performance have focused on undamaged structures without any signs of deterioration. This paper focuses on the analysis of the effects of the initial damage state on the seismic response and fragility of a five-story CLT building designed under a force-based approach. A detailed 3D finite element model was developed and validated through experimental data in order to perform incremental dynamic analyses that considered different arbitrarily imposed initial damage states. The residual response and the fragility functions are analyzed to characterize the impact of the initial state on seismic behavior. The results of this work highlight the need to properly consider the effect of previous load actions for the seismic performance evaluation during the operating life of CLT structures. Findings suggest that the initial state can significantly modify the probability of reaching a given limit state. Moreover, it was found that if the initial damage is defined as severe, the collapse margin ratio is reduced by 58.8% compared to the case in which the initial state is undamaged. Full article
(This article belongs to the Special Issue Research on Timber and Timber–Concrete Buildings)
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<p>Floor plan and elevation views (<b>top</b>). Image during the construction stage, and the finished state of the building (<b>bottom</b>).</p>
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<p>Cross-laminated timber (CLT) shear wall components (<b>right</b>) and model implementation (<b>left</b>).</p>
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<p>Deformation mechanism (<b>top</b>), and lateral load-displacement response comparisons (<b>bottom</b>).</p>
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<p>Two-step analysis scheme developed for the initial damage state assessment.</p>
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<p>Unscaled elastic ground motion spectra of the seismic demands used for the IDA.</p>
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<p>Studied building capacity curves.</p>
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<p>Free vibration (FV) response after the imposed pushover roof displacement. Initial FV response (<b>left</b>) and residual drift at the end of the FV stage (<b>right</b>).</p>
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<p>Residual drift with respect to the imposed roof displacement demand.</p>
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<p>Fundamental frequency degradation (<b>above</b>) and Final Softening Index (DF) evolution (<b>bellow</b>) with respect to roof displacement demand.</p>
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<p>Variation of the median IM with respect to the inter-story drift (DM) and the initial damage level for the X-direction (<b>a</b>) and Y-direction (<b>b</b>). Light gray lines correspond to the IDA curves of each considered seismic demand for all the initial damage states.</p>
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<p>Fragility curves for the following limit states and directions: (<b>a</b>) Fully Operational in X-direction, (<b>b</b>) Life Safety in X-direction, (<b>c</b>) Collapse in Y-direction, (<b>d</b>) Fully Operational in Y-direction, (<b>e</b>) Life Safety in Y-direction, and (<b>f</b>) Collapse in Y-direction.</p>
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<p>Relationship between initial damage residual roof drift (<math display="inline"><semantics> <msub> <mo>Δ</mo> <mi>res</mi> </msub> </semantics></math>) and the Collapse Margin Ratio (CMR).</p>
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17 pages, 8581 KiB  
Article
Enhanced Control Strategy for Three-Level T-Type Converters in Hybrid Power-to-X Systems
by Moria Sassonker Elkayam and Dmitri Vinnikov
Appl. Sci. 2025, 15(5), 2409; https://doi.org/10.3390/app15052409 - 24 Feb 2025
Viewed by 186
Abstract
This paper presents a dual-loop control system designed for three-level three-phase T-type converters, optimizing their performance in the hybrid operation of Power-to-X systems. Due to the increasing of distributed power generation based on renewable energy sources, Power-to-X systems convert surplus renewable energy into [...] Read more.
This paper presents a dual-loop control system designed for three-level three-phase T-type converters, optimizing their performance in the hybrid operation of Power-to-X systems. Due to the increasing of distributed power generation based on renewable energy sources, Power-to-X systems convert surplus renewable energy into other forms of energy, such as hydrogen, synthetic fuels, or chemical storage, which can be stored and later converted back to electricity or used in other applications. Bidirectional converters play a crucial role in hybrid system operation, which requires an efficient and reliable power conversion to maintain stability and performance. The proposed dual-loop control system includes an inner current loop for fast current regulation and an outer voltage loop to maintain stable voltage levels, ensuring precise control of the output of the converter and enhancing its response to dynamic changes in load and generation. Additionally, the control system incorporates a technique to balance the split DC-link capacitors voltages, a major challenge in three-level converters. Comprehensive simulation and experimental results demonstrate the efficacy of the proposed control system in maintaining high power quality and supporting the hybrid operation of Power-to-X systems. Full article
(This article belongs to the Special Issue Control of Power Systems II)
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<p>Block diagram of the proposed hybrid system.</p>
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<p>Bidirectional three-level three-phase T-type converter.</p>
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<p>T-type converter employed in the hybrid system.</p>
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<p>Three-level three-phase T-type converter with the <span class="html-italic">LCL</span> filter.</p>
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<p>Single-phase <span class="html-italic">LCL</span>-filter-based converter representation.</p>
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<p>Block diagram of the proposed control structure.</p>
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<p>(<b>a</b>) Open-loop and (<b>b</b>) closed-loop Bode diagrams of the current loop.</p>
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<p>Equivalent block diagram of the voltage loop.</p>
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<p>Block diagram of the DC-link capacitors voltage balancing control.</p>
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<p>Simulation results of the capacitor voltage reference and output signals of both <span class="html-italic">α</span> (<b>top</b>) and <span class="html-italic">β</span> (<b>bottom</b>) loops.</p>
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<p>Simulation results of the inverter side inductor current reference and output signals of both <span class="html-italic">α</span> (top) and <span class="html-italic">β</span> (bottom) loops.</p>
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<p>Simulation results of the three-phase signals of output voltages (<b>top</b>) and inverter side inductors currents (<b>bottom</b>).</p>
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<p>Simulation results of (<b>a</b>) DC-link capacitors voltages and (<b>b</b>) modulation signals of the three phases before (<b>top</b>) and after (<b>bottom</b>) the correction.</p>
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<p>Simulation results three phase signals of output voltages and inverter side inductors currents.</p>
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<p>A three-level, three-phase T-type <span class="html-italic">LCL</span> filtered converter experimental platform.</p>
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<p>(<b>a</b>) Experiment values of the three phases and (<b>b</b>) experimental results of the DC-link capacitors voltages (<b>top</b> and <b>bottom</b>), the capacitor voltage of phase <span class="html-italic">a</span>, and the inverter current of phase <span class="html-italic">b</span> (The displayed frequency is 49.99 in all phases).</p>
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<p>Comparison between experimental results (<b>left</b>) and simulation (<b>right</b>), including DC side capacitors voltages, the output voltage across the filter capacitor of phase <span class="html-italic">a</span>, and the current of phase <span class="html-italic">b</span>.</p>
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<p>Comparison between experimental results of DC-link capacitors voltages (<b>top</b> and <b>bottom</b>), capacitor voltage of phase <span class="html-italic">a</span>, and inverter current of phase <span class="html-italic">b</span> with PI controller.</p>
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13 pages, 611 KiB  
Review
Permanent Maxillary Lateral Incisors’ Agenesis Managed by Mini-Screw Implant-Supported Pontics: A Scoping Review
by Elena Caramaschi, Elisabetta Lalli, Valentino Garau, Alessio Verdecchia and Enrico Spinas
Dent. J. 2025, 13(3), 96; https://doi.org/10.3390/dj13030096 - 24 Feb 2025
Viewed by 88
Abstract
Background/Objectives: The Agenesis of maxillary lateral incisors occurs with a variable prevalence in different ethnic groups, and there is a need for a temporary replacement until maturity has been reached in patients for whom the replacement solution has been chosen. This study aims [...] Read more.
Background/Objectives: The Agenesis of maxillary lateral incisors occurs with a variable prevalence in different ethnic groups, and there is a need for a temporary replacement until maturity has been reached in patients for whom the replacement solution has been chosen. This study aims to analyze the scientific evidence available to date concerning the use of mini-screw implant (MSI)-supported pontics for the transitional management of permanent maxillary lateral incisors’ agenesis in developmental age subjects. Methods: Electronic research was conducted using four databases: PubMed, Clarivate Analytics/Web of Science Core Collection, Scopus, and the Wiley/Cochrane Library. Six studies were included in the final review. Data were extracted based on the first and second author, year of publication, study design, sample characteristics, mini-screw implant (MSI) characteristics, MSI insertion and loading protocol, characteristics of the prosthetic component, and outcomes during the follow-up time. Results: Clinical outcomes were proven positive in all studies. In only one study did MSIs show mobility and consequent failure after one month. Discoloration of the prosthetic part proved to be the main complication. Conclusions: The comparison with conventional removable prostheses and fixed dental prostheses revealed that MSI-supported pontics are a viable alternative and a promising temporary solution until the end of growth. Further studies to standardize protocols and assess long-term outcomes are needed. Full article
(This article belongs to the Special Issue Dentistry in the 21st Century: Challenges and Opportunities)
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Graphical abstract

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<p>Flow chart of the search strategy and selection of articles to be included.</p>
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