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19 pages, 6946 KiB  
Article
Influence of Particle Rotation on the Shear Characteristics of Calcareous-Sand and Silica-Bead Granular Materials
by Tao Li, Jiajun Shu, Yuming Wu, Yue Li, Bingni Wu, Zhengding Deng, Jingzhu Huang, Rubén Galindo and Fausto Molina Gómez
Materials 2024, 17(23), 5827; https://doi.org/10.3390/ma17235827 (registering DOI) - 27 Nov 2024
Abstract
The shear strength and resistance of granular materials are critical indicators in geotechnical engineering and infrastructure construction. Both sliding and rotation influence the energy evolution of soil granular motion during shear. To examine the effects of particle rotation on shear damage and energy [...] Read more.
The shear strength and resistance of granular materials are critical indicators in geotechnical engineering and infrastructure construction. Both sliding and rotation influence the energy evolution of soil granular motion during shear. To examine the effects of particle rotation on shear damage and energy evolution in granular systems, we first describe the transformation of irregularly shaped particles into regular shapes via geometrical parameters, ensuring the invariance of energy density and density. We then analyze the impact of particle rotation on shear-stress variation and energy dissipation through a shear energy evolution equation. Additionally, we establish the relationship between the shear-stress ratio and normal stress, considering particle rotation. Finally, we verify the influence of particle rotation on energy evolution and shear damage through shear tests on irregular calcareous sand and regular silica-bead particles. The results indicate that granular materials do not fully comply with the Coulomb strength criterion. In the initial shear stage, most of the external work is converted into granular rotational-shear energy, whereas in the later stage, it primarily shifts to granular sliding-shear energy. Notably, the sensitivity of the granular rotational energy to a vertical load is significantly greater than that of the granular sliding energy. Full article
(This article belongs to the Special Issue Preparation and Application of Regularly Structured Porous Materials)
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<p>Schematic characterization of the shape parameters of granular materials.</p>
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<p>Electron microscopy images of the two granular materials.</p>
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<p>Schematic diagram of the elliptical wrapping method.</p>
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<p>Schematic of particle rotation.</p>
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<p>Relationship between the shear-stress-to-positive-stress ratio and the shear displacement of glass beads.</p>
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<p>Relationship between the shear-stress-to-positive-stress ratio and the shear displacement of calcareous sand.</p>
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<p>Relationships between the particle rotational properties and the shear displacement of glass beads.</p>
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<p>Relationships between the particle rotational properties and the shear displacement of calcareous sand.</p>
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<p>Relationship between the angle of rotation and the shear displacement of glass beads.</p>
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<p>Relationship between the angle of rotation and the shear displacement of calcareous sand.</p>
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<p>Strain-energy density evolution in glass-bead granular systems (under 100 kPa).</p>
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<p>Strain-energy density evolution in calcareous-sand granular systems (under 100 kPa).</p>
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<p>Strain-energy density evolution in glass-bead granular systems (under 200 kPa).</p>
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<p>Strain-energy density evolution in calcareous-sand granular systems (under 200 kPa).</p>
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<p>Strain-energy density evolution in glass-bead granular systems (under 300 kPa).</p>
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<p>Strain-energy density evolution in calcareous-sand granular systems (under 300 kPa).</p>
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<p>Sources and percentages of the shear strengths of the glass-bead granular materials (after averaging).</p>
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<p>Sources and percentages of the shear strengths of the calcareous-sand granular materials (after averaging).</p>
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19 pages, 3753 KiB  
Article
Seismic Resistance of Reinforced Concrete Building Frames Based on Interval Assessment of the Coefficient of Permissible Damage
by Ashot Tamrazyan and Tatiana Matseevich
Buildings 2024, 14(12), 3776; https://doi.org/10.3390/buildings14123776 - 26 Nov 2024
Viewed by 191
Abstract
The main method for assessing the seismic resistance of buildings in the standards of most countries is the linear-spectral method. This method allows for the calculation of the spatial model of a building for seismic load in the elastic range without resorting to [...] Read more.
The main method for assessing the seismic resistance of buildings in the standards of most countries is the linear-spectral method. This method allows for the calculation of the spatial model of a building for seismic load in the elastic range without resorting to direct integration of the equations of motion. Nonlinear characteristics of reinforced concrete structure materials are usually considered integrally using the reduction factor. However, the values of this factor in the Russian standards are not sufficiently substantiated, as the later studies show. To determine the coefficient of permissible damage (reduction factor), six reinforced concrete frames were considered, with different parameters such as span length, number of spans, and number of floors. The design parameters of beams and columns (section sizes, reinforcement, etc.) were preliminarily selected based on the calculation using the linear-spectral method. In the second stage, numerical modeling was carried out in the OpenSEES PC to implement the pushover analysis procedure. Then, the coefficient of permissible damage was estimated by processing the capacity curves obtained on the basis of nonlinear static calculation. The value of the sought coefficients is practically not affected by the number of stores of the frame; however, with an increase in the number of spans, the coefficient K1 increases, which is explained by a decrease in the plasticity of the system. On average, for the frames under consideration, the coefficient K1 was 0.526, which is 1.5 times greater than the coefficient proposed in modern Russian standards, K1 = 0.35. The results obtained on the basis of pushover analysis are compared with the coefficients K1 determined through the values of the average degree of damage (d) of the buildings according to the modified seismic scale MMSK-86. For various types of reinforced concrete frame buildings, K1 = 0.51 was obtained. It is recommended that the coefficient K1 for reinforced concrete frame buildings should be increased to a value of at least K1 = 0.5 in the Russian standard. Full article
(This article belongs to the Special Issue Safety and Optimization of Building Structures—2nd Edition)
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<p>To determine the reduction coefficient: 1—linear elastic system; 2—nonlinear system.</p>
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<p>The reinforced concrete frames under consideration.</p>
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<p>Calculation scheme of reinforced concrete frame P-6-3-5 in PC OpenSEES.</p>
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<p>Calculated response spectrum.</p>
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<p>Material deformation diagrams.</p>
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<p>Horizontal loads on the floors of the frames with (<b>a</b>) 5 and (<b>b</b>) 7 stories.</p>
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<p>To determine the coefficient of permissible damage.</p>
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<p>Reinforcement scheme of frame elements.</p>
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<p>The distribution of chord rotations.</p>
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<p>Bearing capacity curves of reinforced concrete frames.</p>
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<p>Calculation example for frame P-6-3-5.</p>
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<p>Graph of dependence of <span class="html-italic">K</span><sub>1</sub> on the average degree of damage to buildings <span class="html-italic">d</span>.</p>
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18 pages, 4211 KiB  
Article
Effect of Traffic Vibration on Compressive Strength of High-Strength Concrete and Tensile Strength of New-to-Old Concrete Interfaces
by Pingping Gu, Hao Wu, Luchang Li, Zhanghao Li, Jingyi Hong and Mei-Ling Zhuang
Buildings 2024, 14(12), 3765; https://doi.org/10.3390/buildings14123765 - 26 Nov 2024
Viewed by 229
Abstract
Widening existing bridges is an important way to meet the surge in traffic demand, which is often carried out in a way that does not interrupt traffic. To investigate the effect of traffic vibration on the compressive strength of high-strength concrete and the [...] Read more.
Widening existing bridges is an important way to meet the surge in traffic demand, which is often carried out in a way that does not interrupt traffic. To investigate the effect of traffic vibration on the compressive strength of high-strength concrete and the splitting strength of new-to-old concrete interfaces, the initial to final set time of high-strength concrete C60 was first investigated in this article. Then, the traffic disturbance parameters were determined. Later, the compressive strength of C60 concrete at different stages under traffic disturbance parameters was carried out. Finally, the splitting tensile strength of new-to-old concrete specimens at different stages with different loading modes was tested. The test results indicated that the compressive strength of the specimens vibrated for 3 h and cured for 3, 7, and 28 days was increased by 4.3%, 5.7%, and 11.9%, respectively; those of the specimens vibrated for 7 h and cured for 3, 7, and 28 days was decreased by 13.7%, 20.4%, and 19.9%, respectively; the effect traffic vibration on the compressive strength of the specimens vibrated for 5 h was not obvious. When loaded along the old and new concrete joint, the specimens cracked along the joint; the splitting tensile strengths of the specimen at different disturbed stages were significantly decreased. When loaded perpendicular to the joint, the specimens cured for 3 and 7 days still cracked along the joint, and the splitting tensile strengths of the specimen at different disturbed stages were significantly decreased; while the specimens cured for 28 days cracked in the direction perpendicular to the joint, the tensile strengths of the specimens at different disturbed stages were significantly decreased. This study can promote the widening and improvement of existing concrete highways and bridges, which can save resources and improve land use. Full article
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<p>Test molds and a penetration resistance tester.</p>
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<p>Cross-section of the widening bridge (Unit: mm) (The blue in <a href="#buildings-14-03765-f002" class="html-fig">Figure 2</a> for the new concrete bridge; the red in <a href="#buildings-14-03765-f002" class="html-fig">Figure 2</a> for the new-to-old interface).</p>
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<p>Data acquisition equipment.</p>
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<p>On-site testing.</p>
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<p>Vertical vibration testbed.</p>
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<p>Cube compressive strength testing device.</p>
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<p>New-to-old concrete specimens.</p>
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<p>Cube splitting tensile strength test.</p>
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<p>Concrete failure modes.</p>
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<p>Compressive strength of concrete under travel disturbance.</p>
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<p>Damage morphology.</p>
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<p>Comparison of splitting tensile strengths of specimens loaded along the direction of the joints.</p>
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<p>Comparison of splitting tensile strength of specimens loaded along the direction perpendicular to the joints.</p>
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23 pages, 9192 KiB  
Article
Seismic Behavior of Resilient Reinforced Concrete Columns with Ultra-High-Strength Rebars Under Strong Earthquake-Induced Multiple Reversed Cyclic Loading
by Yue Wen, Gaochuang Cai, Prafulla Bahadur Malla, Hayato Kikuchi and Cheng Xie
Buildings 2024, 14(12), 3747; https://doi.org/10.3390/buildings14123747 - 25 Nov 2024
Viewed by 271
Abstract
The frequent occurrence of major earthquakes highlights the structural challenges posed by long-period ground motions (LPGMs). This study investigates the seismic performance and resilience of five reinforced concrete (RC) columns with different high-strength steel bars under LPGM-induced cyclic loading, both experimentally and numerically. [...] Read more.
The frequent occurrence of major earthquakes highlights the structural challenges posed by long-period ground motions (LPGMs). This study investigates the seismic performance and resilience of five reinforced concrete (RC) columns with different high-strength steel bars under LPGM-induced cyclic loading, both experimentally and numerically. The results show that low-bond and debonded high-strength steel bars significantly enhance self-centering capabilities and reduce residual drift, with lateral force reductions of 7.6% for normal cyclic loading and 19.2% for multiple reversed cyclic loading. The concrete damage in the hinge zone of the columns was increased; however, the significant inside damage of the concrete near the steel bars made it easier to restore the columns for the damage accumulation caused by multiple loading. Based on the experiment, a numerical model was developed for the columns, and a simplified model was proposed to predict energy dissipation capacity, providing practical design methods for resilient RC structures that may be attacked by LPGMs. Full article
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<p>Details of tested specimens: (<b>a</b>) F60S3U; (<b>b</b>) F60S3SB and F60S3SBD.</p>
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<p>Used reinforcements in the study: (<b>a</b>) SBPDN1275/1420 bar (U); (<b>b</b>) SBPDN1080/1230 bar (USD).</p>
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<p>Test methods: (<b>a</b>) test setup; (<b>b</b>) LVDT position.</p>
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<p>Loading protocols.</p>
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<p>Damage development and failure of the tested columns (Blue: pull direction, Red: push direction).</p>
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<p>(<b>a</b>) F60S3U-NC; (<b>b</b>) F60S3SB-NC; (<b>c</b>) F60S3SB-MRC; (<b>d</b>) F60S3SBD-NC; (<b>e</b>) F60S3SBD-MRC.</p>
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<p>Skeleton curves of the specimens: (<b>a</b>) F60S3U-NC; (<b>b</b>) comparison between F60S3SB-NC and F60S3SB-MRC; (<b>c</b>) comparison between F60S3SBD-NC and F60S3SBD-MRC.</p>
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<p>Comparison of the envelope curves of the specimens: (<b>a</b>) effect of loading method; (<b>b</b>) effect of unbonded length.</p>
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<p>Residual drift ratios of the specimens: (<b>a</b>) specimens reinforced by different rebars; (<b>b</b>) specimens reinforced by USD and unbonded USD rebars.</p>
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<p>Calculation method of equivalent viscous damping coefficient.</p>
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<p>Equivalent viscous damping coefficients of tested specimens.</p>
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<p>Column modeling and analysis method.</p>
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<p>Stress–strain model of concrete.</p>
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<p>Stress–strain model of steel rebar.</p>
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<p>Bond–slip models of steel bars: (<b>a</b>) Funato et al.’s model for U-steel bars [<a href="#B26-buildings-14-03747" class="html-bibr">26</a>]; (<b>b</b>) Shima et al.’s model for USD bars [<a href="#B33-buildings-14-03747" class="html-bibr">33</a>].</p>
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<p>(<b>a</b>–<b>c</b>) Comparison of skeleton curve between experimental and simulated results.</p>
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<p>Effects of the compressive strength of concrete: (<b>a</b>) specimens reinforced with low-bond ultra-high-strength rebars (U series); (<b>b</b>) specimens reinforced with ultra-high-strength USD rebars (SB series); (<b>c</b>) specimens reinforced with unbonded ultra-high-strength USD rebars (SBD series).</p>
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<p>Effects of the shear span ratios (a/D): (<b>a</b>) specimens reinforced with low-bond ultra-high-strength rebars (U series); (<b>b</b>) specimens reinforced with ultra-high-strength USD rebars (SB series); (<b>c</b>) specimens reinforced with unbonded ultra-high-strength USD rebars (SBD series).</p>
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<p>Comparison between experimental and simulated curves: (<b>a</b>) SB columns, (<b>b</b>) SBD columns.</p>
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<p>Strength degradation ratio vs. the number of loading cycles.</p>
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<p>Relationship between frequency and degradation at drift ratio 2%.</p>
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<p>Comparison between experimental and adjusted curves: (<b>a</b>) SB columns, (<b>b</b>) SBD columns.</p>
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<p>Proposal of a simplified equivalent viscous damping coefficient model.</p>
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<p>Comparison between the test and analysis results: (<b>a</b>) MRC specimens, (<b>b</b>) NC specimens.</p>
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19 pages, 4517 KiB  
Article
Assessment of the Actual and Required Cooling Demand for Buildings with Extensive Transparent Surfaces
by Attila Kostyák, Szabolcs Szekeres and Imre Csáky
Energies 2024, 17(23), 5814; https://doi.org/10.3390/en17235814 - 21 Nov 2024
Viewed by 212
Abstract
Energy consumption in buildings with large, glazed facades rises markedly in the summer, driven by cooling demands that vary with structural characteristics and external climate conditions. This study is unique in examining daily cooling needs in lightweight and heavyweight constructions, utilizing meteorological data [...] Read more.
Energy consumption in buildings with large, glazed facades rises markedly in the summer, driven by cooling demands that vary with structural characteristics and external climate conditions. This study is unique in examining daily cooling needs in lightweight and heavyweight constructions, utilizing meteorological data from 782 summer days in Debrecen, Hungary. Unlike standard approaches, which often overlook localized meteorological variables, this analysis focuses on actual “clear sky” scenarios across distinct summer day types: normal, hot, and torrid. The findings indicate that orientation and construction type significantly affect cooling demands, with east-facing rooms demanding up to 14.2% more cooling in lightweight structures and up to 35.8% in heavyweight structures during peak hours (8 a.m. to 4 p.m.). This study reveals that for west-facing facades, extending use beyond 4 p.m. markedly increases energy loads. Furthermore, the cooling demand peak for heavyweight buildings occurs later in the day, driven by their higher thermal capacity. These insights underscore the importance of aligning HVAC system design with operational schedules and building orientation, offering data-driven strategies to enhance energy efficiency in buildings with diverse thermal and solar exposure profiles. Full article
(This article belongs to the Special Issue Energy Efficiency of the Buildings III)
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<p>The Great Hungarian Plain [<a href="#B18-energies-17-05814" class="html-bibr">18</a>].</p>
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<p>Daily variation of <span class="html-italic">R<sub>b</sub></span> values for different orientations of the wall on a summer day.</p>
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<p>Direct solar radiation values.</p>
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<p>Comparison of design days by external air temperature.</p>
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<p>Comparison of design days by energy yield.</p>
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<p>Example building.</p>
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<p>RC network heat flows [<a href="#B4-energies-17-05814" class="html-bibr">4</a>].</p>
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<p>(<b>a</b>) Summarized cooling demand. (<b>b</b>–<b>e</b>) Cooling demand of a room with the transparent surface facing (<b>b</b>) north, (<b>c</b>) south, (<b>d</b>) east, and (<b>e</b>) west.</p>
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<p>(<b>a</b>–<b>f</b>) Cooling demand of sample building at different orientations in the case of summer, hot, and torrid days.</p>
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<p>(<b>a</b>–<b>f</b>) Cooling demand of sample building at different orientations in the case of summer, hot, and torrid days.</p>
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22 pages, 9902 KiB  
Article
Analytical Fragility Surfaces and Global Sensitivity Analysis of Buried Operating Steel Pipeline Under Seismic Loading
by Gersena Banushi
Appl. Sci. 2024, 14(22), 10735; https://doi.org/10.3390/app142210735 - 20 Nov 2024
Viewed by 318
Abstract
The structural integrity of buried pipelines is threatened by the effects of Permanent Ground Deformation (PGD), resulting from seismic-induced landslides and lateral spreading due to liquefaction, requiring accurate analysis of the system performance. Analytical fragility functions allow us to estimate the likelihood of [...] Read more.
The structural integrity of buried pipelines is threatened by the effects of Permanent Ground Deformation (PGD), resulting from seismic-induced landslides and lateral spreading due to liquefaction, requiring accurate analysis of the system performance. Analytical fragility functions allow us to estimate the likelihood of seismic damage along the pipeline, supporting design engineers and network operators in prioritizing resource allocation for mitigative or remedial measures in spatially distributed lifeline systems. To efficiently and accurately evaluate the seismic fragility of a buried operating steel pipeline under longitudinal PGD, this study develops a new analytical model, accounting for the asymmetric pipeline behavior in tension and compression under varying operational loads. This validated model is further implemented within a fragility function calculation framework based on the Monte Carlo Simulation (MCS), allowing us to efficiently assess the probability of the pipeline exceeding the performance limit states, conditioned to the PGD demand. The evaluated fragility surfaces showed that the probability of the pipeline exceeding the performance criteria increases for larger soil displacements and lengths, as well as cover depths, because of the greater mobilized soil reaction counteracting the pipeline deformation. The performed Global Sensitivity Analysis (GSA) highlighted the influence of the PGD and soil–pipeline interaction parameters, as well as the effect of the service loads on structural performance, requiring proper consideration in pipeline system modeling and design. Overall, the proposed analytical fragility function calculation framework provides a useful methodology for effectively assessing the performance of operating pipelines under longitudinal PGD, quantifying the effect of the uncertain parameters impacting system response. Full article
(This article belongs to the Section Civil Engineering)
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Graphical abstract

Graphical abstract
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<p>Pipeline subjected to longitudinal PGD: (<b>a</b>) 3D view; (<b>b</b>) 2D schematic representation.</p>
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<p>Pipeline response to longitudinal PGD according to analytical model in [<a href="#B11-applsci-14-10735" class="html-bibr">11</a>], assuming symmetric material behavior for tension and compression: (<b>a</b>) case I; (<b>b</b>) case II.</p>
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<p>Schematic representation of operating pipeline response subjected to longitudinal PGD: (<b>a</b>) pipeline displacement subjected to longitudinal soil block movement (case II); (<b>b</b>) soil–pipeline system behaving like a pull-out test under tension (region I) and compression (region IV).</p>
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<p>Schematic representation of the axial constitutive behavior of the steel pipe material, defined within the associated von Mises plasticity with isotropic hardening [<a href="#B30-applsci-14-10735" class="html-bibr">30</a>].</p>
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<p>The comparison between the numerical, the conventional [<a href="#B8-applsci-14-10735" class="html-bibr">8</a>,<a href="#B11-applsci-14-10735" class="html-bibr">11</a>,<a href="#B13-applsci-14-10735" class="html-bibr">13</a>], and the proposed analytical models, evaluating the pipeline performance under longitudinal PGD (<span class="html-italic">L<sub>b</sub></span> = 300 m) in terms of maximum tensile and compressive pipe strain as a function of the ground displacement <span class="html-italic">δ</span>.</p>
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<p>The variation of the critical soil block length, <span class="html-italic">L<sub>cr</sub></span> = (<span class="html-italic">F<sub>t,max</sub></span> − <span class="html-italic">F<sub>c,max</sub></span>)/<span class="html-italic">f<sub>s</sub></span>, as a function of the ground displacement <span class="html-italic">δ</span>, with an indication of the critical values (<span class="html-italic">δ<sub>cr,i</sub></span>, <span class="html-italic">L<sub>cr,i</sub></span>) associated with the achievement of the pipeline performance limit states.</p>
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<p>The peak axial strain magnitude in the pressurized pipeline (<span class="html-italic">P<sub>i</sub></span>/<span class="html-italic">P<sub>max</sub></span> = 0.75, Δ<span class="html-italic">T</span> = 50 °C) as a function of the PGD length <span class="html-italic">L<sub>b</sub></span> and displacement <span class="html-italic">δ</span> for (<b>a</b>) tension and (<b>b</b>) compression. The dashed horizontal curves represent the strain isolines corresponding to the NOL and PIL performance limit states.</p>
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<p>The peak axial strain magnitude in the unpressurized pipeline (<span class="html-italic">P<sub>i</sub></span>/<span class="html-italic">P<sub>max</sub></span> = 0, Δ<span class="html-italic">T</span> = 0 °C) as a function of the PGD length <span class="html-italic">L<sub>b</sub></span> and displacement <span class="html-italic">δ</span> for (<b>a</b>) tension and (<b>b</b>) compression. The dashed horizontal curves represent the strain isolines corresponding to the NOL and PIL performance limit states.</p>
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<p>Fragility surface of buried pipeline (<span class="html-italic">H<sub>c</sub></span> = 1.5 m) for (<b>a</b>) Normal Operability Limit (NOL) and (<b>b</b>) Pressure Integrity Limit (PIL).</p>
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<p>Schematic representation of the performance assessment of the buried pipeline subjected to the PGD demand (<span class="html-italic">δ</span>, <span class="html-italic">L<sub>b</sub></span>), using the deterministic and fragility analysis framework.</p>
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<p>Fragility surface of buried pipeline for different cover depths and performance limit states: (<b>a</b>) <span class="html-italic">H<sub>c</sub></span> = 1.0 m, NOL; (<b>b</b>) <span class="html-italic">H<sub>c</sub></span> = 1.0 m, PIL and (<b>c</b>) <span class="html-italic">H<sub>c</sub></span> = 2.0 m, NOL; and (<b>d</b>) <span class="html-italic">H<sub>c</sub></span> = 2.0 m, PIL.</p>
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<p>The comparison of the first-order and total-order sensitivity indices of the system input parameters for the (<b>a</b>) NOL and (<b>b</b>) PIL performance limit states.</p>
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<p>Response of the pressurized pipeline (<span class="html-italic">P<sub>i</sub></span>/<span class="html-italic">P<sub>max</sub></span> = 0.75, Δ<span class="html-italic">T</span> = 50 °C) to longitudinal PGD with block length <span class="html-italic">L<sub>b</sub></span> = 200 m (case I): (<b>a</b>) pipe axial force; (<b>b</b>) pipe axial stress; (<b>c</b>) soil friction; (<b>d</b>) ground displacement; (<b>e</b>) pipe axial displacement; (<b>f</b>) pipe axial strain vs. distance from tension crack.</p>
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<p>Response of the unpressurized pipeline (<span class="html-italic">P<sub>i</sub></span>/<span class="html-italic">P<sub>max</sub></span> = 0, Δ<span class="html-italic">T</span> = 0 °C) to longitudinal PGD with block length <span class="html-italic">L<sub>b</sub></span> = 200 m (case I): (<b>a</b>) pipe axial force; (<b>b</b>) pipe axial stress; (<b>c</b>) soil friction; (<b>d</b>) ground displacement; (<b>e</b>) pipe axial displacement; (<b>f</b>) pipe axial strain vs. distance from tension crack.</p>
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<p>Response of the pressurized pipeline (<span class="html-italic">P<sub>i</sub></span>/<span class="html-italic">P<sub>max</sub></span> = 0.75, Δ<span class="html-italic">T</span> = 50 °C) to longitudinal PGD with block length <span class="html-italic">L<sub>b</sub></span> = 300 m (case II): (<b>a</b>) pipe axial force; (<b>b</b>) pipe axial stress; (<b>c</b>) soil friction; (<b>d</b>) ground displacement; (<b>e</b>) pipe axial displacement; (<b>f</b>) pipe axial strain vs. distance from tension crack.</p>
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<p>Response of the pressurized pipeline (<span class="html-italic">P<sub>i</sub></span>/<span class="html-italic">P<sub>max</sub></span> = 0, Δ<span class="html-italic">T</span> = 0 °C) to longitudinal PGD with block length <span class="html-italic">L<sub>b</sub></span> = 300 m (case II): (<b>a</b>) pipe axial force; (<b>b</b>) pipe axial stress; (<b>c</b>) soil friction; (<b>d</b>) ground displacement; (<b>e</b>) pipe axial displacement; (<b>f</b>) pipe axial strain vs. distance from tension crack.</p>
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15 pages, 7825 KiB  
Technical Note
D-InSAR-Based Analysis of Slip Distribution and Coulomb Stress Implications from the 2024 Mw 7.01 Wushi Earthquake
by Yurong Ding, Xin Liu, Xiaofeng Dai, Gaoying Yin, Yang Yang and Jinyun Guo
Remote Sens. 2024, 16(22), 4319; https://doi.org/10.3390/rs16224319 - 19 Nov 2024
Viewed by 255
Abstract
On 23 January 2024, an Mw 7.01 earthquake struck the Wushi County, Xinjiang Uygur Autonomous Region, China. The occurrence of this earthquake provides an opportunity to gain a deeper understanding of the rupture behavior and tectonic activity of the fault system in [...] Read more.
On 23 January 2024, an Mw 7.01 earthquake struck the Wushi County, Xinjiang Uygur Autonomous Region, China. The occurrence of this earthquake provides an opportunity to gain a deeper understanding of the rupture behavior and tectonic activity of the fault system in the Tianshan seismic belt. The coseismic deformation field of the Wushi earthquake was derived from Sentinel-1A ascending and descending track data using Differential Interferometric Synthetic Aperture Radar (D-InSAR) technology. The findings reveal a maximum line-of-sight (LOS) displacement of 81.1 cm in the uplift direction and 16 cm in subsidence. Source parameters were determined using an elastic half-space dislocation model. The slip distribution on the fault plane for the Mw 7.01 Wushi earthquake was further refined through a coseismic slip model, and Coulomb stress changes on nearby faults were calculated to evaluate seismic hazards in surrounding areas. Results indicate that the coseismic rupture in the Mw 7.01 Wushi earthquake sequence was mainly characterized by left-lateral strike-slip motion. The peak fault slip was 3.2 m, with a strike of 228.34° and a dip of 61.80°, concentrated primarily at depths between 5 and 25 km. The focal depth is 13 km. This is consistent with findings reported by organizations like the United States Geological Survey (USGS). The fault rupture extended to the surface, consistent with field investigations by the Xinjiang Uygur Autonomous Region Earthquake Bureau. Coulomb stress results suggest that several fault zones, including the Kuokesale, Dashixia, Piqiang North, Karaitike, southeastern sections of the Wensu, northwestern sections of the Tuoergan, and the Maidan-Sayram Fault Zone, are within regions of stress loading. These areas show an increased risk of future seismic activity and warrant close monitoring. Full article
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<p>Major Cenozoic faults and tectonic background of the 2024 M<sub>w</sub> 7.01 Wushi earthquake in the Tianshan Belt. White arrows represent the GPS horizontal velocity field; black jagged lines indicate faults; yellow stars mark the locations of earthquakes of magnitude 6 or above since 1900 (<a href="#app1-remotesensing-16-04319" class="html-app">Table S1</a>) (USGS: <a href="https://earthquake.usgs.gov/earthquakes/" target="_blank">https://earthquake.usgs.gov/earthquakes/</a>, accessed on 23 May 2024); the red star represents the epicenter; the beach ball symbol indicates the focal mechanism solution of this earthquake (USGS, accessed on 23 May 2024); the lower left corner shows a broader area, and the red box highlights the specific study area; and orange circles represent aftershocks. Fault abbreviations are as follows: TSF, Tuoergan Fault; KKSF, Kuokesale Fault; MSF, Maidan-Sayram Fault; NWSF, North Wensu Fault; KF, Keping Fault; YF, Yimugantawu Fault; KTF, Karatake Fault; and DSXF, Dashixia Fault. The fault data come from the National Earthquake Science Data Center (<a href="https://data.earthquake.cn/" target="_blank">https://data.earthquake.cn/</a>, accessed on 23 May 2024) and the topographic data used are the ETOPO1 Global Relief Model (<a href="https://www.ncei.noaa.gov/products/etopo-global-relief-model" target="_blank">https://www.ncei.noaa.gov/products/etopo-global-relief-model</a>, accessed on 13 August 2024).</p>
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<p>Ground area covered by SAR images utilized for the M<sub>w</sub> 7.01 Wushi earthquake analysis. The red star is the epicenter of the M<sub>w</sub> 7.01 Wushi earthquake; the beach ball represents the focal mechanism solution (USGS, accessed on 23 May 2024); the black lines are faults.</p>
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<p>The LOS deformation of InSAR observations; (<b>a</b>–<b>c</b>) represent the LOS deformation results of track 56, track 34, and track 136, respectively. The yellow star marks the earthquake’s epicenter, and the beach ball symbol represents the focal mechanism solution. The black lines are faults.</p>
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<p>The posterior probability distribution of the fault model parameters for the M<sub>w</sub> 7.01 Wushi earthquake. The red line represents the maximum posterior probability solution. Scatter plots are contoured according to frequency (cold colors for low frequency, warm colors for high frequency).</p>
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<p>Observed Coseismic Interferograms, Modeled Interferograms of the 2024 M<sub>w</sub> 7.01 Wushi Earthquake, and the Residuals After Subtracting the Modeled Deformation from the Observed Interferograms: (<b>a</b>–<b>c</b>) for Track 56, (<b>d</b>–<b>f</b>) for Track 34, and (<b>g</b>–<b>i</b>) for Track 136. The black lines represent faults. The red star indicates the epicenter.</p>
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<p>Trade-off curve showing the balance between slip model roughness and data fit, with the red dot marking the optimal slip factor.</p>
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<p>Coseismic slip distribution for the seismogenic fault of the M<sub>w</sub> 7.01 Wushi earthquake. Arrows are slip vectors as obtained from the inversion.</p>
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<p>The 3D Coseismic slip distribution for the seismogenic fault of the M<sub>w</sub> 7.01 Wushi earthquake. The black dots represent aftershocks.</p>
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<p>Static ΔCFS in neighboring regions induced by the 2024 M<sub>w</sub> 7.01 Wushi earthquake. The black lines are the faults, the red beach ball is the focal mechanism solution, and the yellow five-pointed star is the epicenter.</p>
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<p>Static ΔCFS and aftershock distribution in neighboring regions induced by the 2024 M<sub>w</sub> 7.01 Wushi earthquake. The black lines are the faults, the red beach ball is the focal mechanism solution, and the yellow five-pointed star is the epicenter. White dots are aftershocks.</p>
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22 pages, 14659 KiB  
Article
Effect of Relative Density on the Lateral Response of Piled Raft Foundation: An Experimental Study
by Mohammad Ilyas Siddiqi, Hamza Ahmad Qureshi, Irfan Jamil and Fahad Alshawmar
Buildings 2024, 14(11), 3687; https://doi.org/10.3390/buildings14113687 - 19 Nov 2024
Viewed by 409
Abstract
The population surge has led to a corresponding increase in the demand for high-rise buildings, bridges, and other heavy structures. In addition to gravity loads, these structures must withstand lateral loads from earthquakes, wind, ships, vehicles, etc. A piled raft foundation (PRF) has [...] Read more.
The population surge has led to a corresponding increase in the demand for high-rise buildings, bridges, and other heavy structures. In addition to gravity loads, these structures must withstand lateral loads from earthquakes, wind, ships, vehicles, etc. A piled raft foundation (PRF) has emerged as the most favored system for high-rise buildings due to its ability to resist lateral loads. An experimental study was conducted on three different piled raft model configurations with three different relative densities (Dr) to determine the effect of Dr on the lateral response of a PRF. A model raft was constructed using a 25 mm thick aluminum plate with dimensions of 304.8 mm × 304.8 mm, and galvanized iron (GI) pipes, each 457.2 mm in length, were used to represent the piles. The lateral and vertical load cells were connected to measure the applied loads. It was found that an increase in Dr increased the soil stiffness and led to a decrease in the lateral displacement for all three PRF models. Additionally, the contribution of the piles in resisting the lateral load decreased, whereas the contribution of the raft portion in resisting the lateral load increased. With an increase in Dr from 30% to 90%, the percentage contribution of the raft increased from 42% to 66% for 2PRF, 38% to 61% for 4PRF, and 46% to 70% for 6PRF. Full article
(This article belongs to the Special Issue Advances in Foundation Engineering for Building Structures)
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<p>Particle size distribution of the soil used in this study.</p>
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<p>Direct shear test results for D<sub>r</sub> 30%.</p>
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<p>Direct shear test results for Dr 60%.</p>
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<p>Direct shear test results for Dr 90%.</p>
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<p>(<b>a</b>) Model soil box; (<b>b</b>) horizontal and diagonal stiffeners; (<b>c</b>) 3D view of the box.</p>
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<p>(<b>a</b>) Model raft; (<b>b</b>) raft with hook.</p>
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<p>Model piles.</p>
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<p>Top view schematic diagram of: (<b>a</b>) 2PRF; (<b>b</b>) 4PRF; (<b>c</b>) 6PRF.</p>
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<p>Real PRF models: (<b>a</b>) 2PRF; (<b>b</b>) 4PRF; (<b>c</b>) 6PRF.</p>
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<p>Strain gauge installation.</p>
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<p>Strain gauge calibration process. (<b>a</b>) Digital balance. (<b>b</b>) Load arrangement for calibration.</p>
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<p>Variation in relative density with free fall height.</p>
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<p>(<b>a</b>) Verification of D<sub>r</sub> with DCP; (<b>b</b>) positions for performing DCP.</p>
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<p>Comparison curve for DCP results.</p>
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<p>Schematic diagram of test setup.</p>
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<p>(<b>a</b>) Vertical load cell; (<b>b</b>) lateral load cell; (<b>c</b>) LVDTs.</p>
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<p>Vertical load setup.</p>
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<p>Lateral load setup.</p>
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<p>Analysis of 2PRF at D<sub>r</sub> 30%.</p>
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<p>Analysis of 2PRF at D<sub>r</sub> 60%.</p>
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<p>Analysis of 2PRF at D<sub>r</sub> 90%.</p>
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<p>Analysis of 4PRF at D<sub>r</sub> 30%.</p>
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<p>Analysis of 4PRF at D<sub>r</sub> 60%.</p>
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<p>Analysis of 4PRF at D<sub>r</sub> 90%.</p>
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<p>Analysis of 6PRF at D<sub>r</sub> 30%.</p>
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<p>Analysis of 6PRF at D<sub>r</sub> 60%.</p>
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<p>Analysis of 6PRF at Dr 90%.</p>
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<p>Percentage contribution of 2PRF.</p>
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<p>Percentage contribution of 4PRF.</p>
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<p>Percentage contribution of 6PRF.</p>
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27 pages, 6162 KiB  
Article
Numerical Study of the Nonlinear Soil–Pile–Structure Interaction Effects on the Lateral Response of Marine Jetties
by Marios Koronides, Constantine Michailides, Panagiotis Stylianidis and Toula Onoufriou
J. Mar. Sci. Eng. 2024, 12(11), 2075; https://doi.org/10.3390/jmse12112075 - 17 Nov 2024
Viewed by 366
Abstract
This study presents three-dimensional finite element analyses of two marine structures subjected to lateral loading to approximate environmental forces (e.g., wind, waves, currents, earthquakes). The first structure is a marine jetty supported by twenty-four piles, representative of an existing structure in Cyprus, while [...] Read more.
This study presents three-dimensional finite element analyses of two marine structures subjected to lateral loading to approximate environmental forces (e.g., wind, waves, currents, earthquakes). The first structure is a marine jetty supported by twenty-four piles, representative of an existing structure in Cyprus, while the second is a simplified four-pile marine structure. Soil–pile interaction is modelled using nonlinear p-y, τ-z, and q-z springs that are distributed along the piles, while steel plasticity is also considered. This study examines the relationship between failure modes, deformation modes, and plastic hinge locations with soil behaviour and soil reaction forces. It also aims at investigating the behaviour of the above structures in lateral loading and quantifying the consequences of unrealistic assumptions such as soil and steel linearity or tension-resistant q-z springs. The results indicate that such assumptions can lead to the wrong prediction of failure modes, plastic hinges, and critical elements while emphasising the crucial role of soil nonlinearity and axial pile–soil behaviour on the structural response. It is demonstrated that the dominant nonlinear sources relevant to this study, whether soil nonlinearity, plastic hinge formation, or a combination of the two, are primarily influenced by the axial capacity of soil–pile foundation systems, particularly their tensile component. Full article
(This article belongs to the Section Ocean Engineering)
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<p>Reference marine structure situated off the coast of Vasiliko, Cyprus: (<b>a</b>) view of the entire jetty, (<b>b</b>) T-junction’s closer view.</p>
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<p>View of the platform deck’s underside, depicting the connections between the piles and the deck. It also includes a schematic representation of the pile positions and their inclination directions, as well as details regarding their cross-sectional area and length.</p>
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<p>Soil stratigraphy and material characterisation below the T-junction.</p>
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<p>Profiles of small-strain Young’s modulus derived using Equations (2) and (3), as proposed by [<a href="#B42-jmse-12-02075" class="html-bibr">42</a>,<a href="#B46-jmse-12-02075" class="html-bibr">46</a>], respectively, with the assumed profile superimposed.</p>
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<p>FE model of the SPSI<sup>jetty</sup> system shown in (<b>a</b>) isoparametric, (<b>b</b>) plan (x-y), (<b>c</b>) x-z side, and (<b>d</b>) y-z side views.</p>
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<p>FE model of the SPSI<sup>8×8</sup> system shown in (<b>a</b>) isoparametric, (<b>b</b>) x-z side, and (<b>c</b>) plan (x-y) views.</p>
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<p>Stress–strain behaviour of steel input in the analyses.</p>
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<p>Impact of steel plasticity and nonlinear behaviour of springs on the force–displacement response of SPSI<sup>8×8</sup> marine structure. The stages of plastic hinge formation are illustrated for analyses involving elastoplastic steel.</p>
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<p>Sum of T, Q and P reactions forces of nonlinear springs acting on the −x and +x piles, computed from analyses involving either elastic or elastoplastic steel.</p>
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<p>Sum of T, Q and P reactions forces of linear springs acting on the −x and +x piles, computed from analyses involving either elastic or elastoplastic steel.</p>
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<p>Sequence of plastic hinge formation (indicated by numbering) in the SPSI<sup>8×8</sup> marine structure, as predicted by analyses involving (<b>a</b>) linear springs, and (<b>b</b>) nonlinear springs. Distribution of plastic strains is plotted at the last converged increment of the analyses.</p>
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<p>Impact of steel plasticity, springs nonlinearity and tension allowance of the pile tip springs on the force–displacement response of the SPSI<sup>jetty</sup>.</p>
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<p>Variation in all reaction forces (T, Q and P) of the springs attached on the centre (C) and rear (R) piles, as shown in <a href="#jmse-12-02075-f005" class="html-fig">Figure 5</a>b, with applied lateral force. The results are produced by EPsteel analyses that use either linear or nonlinear springs (nonlinear q-z springs are tensionless).</p>
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<p>Variation in all reaction forces (T, Q and P) of the springs attached on the centre (C) and rear (R) piles, as shown in <a href="#jmse-12-02075-f005" class="html-fig">Figure 5</a>b, with applied lateral force. The results are produced by EPsteel and nonlinear analyses with either tensionless or tension-resistant q-z.</p>
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<p>Plastic strain accumulation on the piles predicted by the analysis with linear springs, illustrated in (<b>a</b>) plan (x-y) view, and (<b>b</b>) side (x-z) view. The numbering indicates the sequence of hinge formation. The side view includes both initial (green) and deformed (black) structures, with displacements exaggerated by a factor of 5. All results are from the last converged increment of the analysis.</p>
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<p>Plastic strain accumulation on the piles predicted by the analysis with nonlinear springs and tensionless q-z springs, illustrated in (<b>a</b>) plan (x-y) view, and (<b>b</b>) side (x-z) view. The numbering indicates the sequence of hinge formation. The side view includes both initial (green) and deformed (black) structures, with displacements exaggerated by a factor of 5. All results are from the last converged increment of the analysis.</p>
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<p>Plastic strain accumulation on the piles predicted by the analysis with nonlinear springs and tension-resistant q-z springs, illustrated in (<b>a</b>) plan (x-y) view, and (<b>b</b>) side (x-z) view. The numbering indicates the sequence of hinge formation. The side view includes both initial (green) and deformed (black) structures, with displacements exaggerated by a factor of 5. All results are from the last converged increment of the analysis.</p>
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<p>(<b>a</b>) Axial force and (<b>b</b>) bending moments acting on various cross-sections (as shown in <a href="#jmse-12-02075-f015" class="html-fig">Figure 15</a>) plotted against the platform’s horizontal displacement.</p>
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<p>T-z spring curves input in the numerical model for (<b>a</b>) dense sand and (<b>b</b>) Marl layers.</p>
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<p>Q-z spring curves input in the numerical model at pile tips.</p>
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<p>P-y spring curves input in the numerical model for (<b>a</b>) dense sand and (<b>b</b>) Marl layers.</p>
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13 pages, 2763 KiB  
Article
Numerical Simulation Study of Bearing Characteristics of Large-Diameter Flexible Piles Under Complex Loads
by Xueying Yang, Weiming Gong and Qian Yin
Buildings 2024, 14(11), 3651; https://doi.org/10.3390/buildings14113651 - 17 Nov 2024
Viewed by 375
Abstract
The majority of the existing calculation methods for determining the ultimate bearing capacity of steel-pipe piles using Chinese criteria are designed for piles with diameters smaller than 2 m. To investigate the bearing capacity of flexible steel-pipe piles with diameters larger than 2 [...] Read more.
The majority of the existing calculation methods for determining the ultimate bearing capacity of steel-pipe piles using Chinese criteria are designed for piles with diameters smaller than 2 m. To investigate the bearing capacity of flexible steel-pipe piles with diameters larger than 2 m under combined loading conditions, reveal nonlinear interactions between vertical and horizontal loads, and propose bearing capacity envelopes, in this paper, a numerical method was used to study the bearing capacity of a flexible pile with a diameter of 2.8 m and an embedment length of 72 m under vertical and horizontal loading conditions. First, a numerical model was developed and calibrated using field test results. Then, the effects of vertical pressure on horizontal capacity and lateral force on vertical capacity and uplift capacity of the pile were analyzed. The results indicate that vertical pressure at the top of the pile can nonlinearly reduce its horizontal capacity, but this pressure initially has a slight positive effect on the horizontal bearing capacity before causing a rapid decrease. Conversely, horizontal force negatively impacts both the compressive and uplift bearing capacities of the pile. Finally, depending on the above results, bearing capacity envelopes for piles subjected to vertical and horizontal loads were proposed. Full article
(This article belongs to the Special Issue Trends and Prospects in Civil Engineering Structures)
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<p>Diagram of the tested pile.</p>
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<p>Diagrams of the numerical simulation: (<b>a</b>) the pile and (<b>b</b>) the soil.</p>
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<p>Comparison of results between tests and simulations: (<b>a</b>) horizontal load–displacement curves and (<b>b</b>) vertical load–displacement curves.</p>
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<p>The horizontal load–displacement curves of a pile at the mudline under different vertical pressures.</p>
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<p>The displacement of a pile along depth under different vertical pressures: (<b>a</b>) <span class="html-italic">V</span> = 1.0 MN, (<b>b</b>) <span class="html-italic">V</span> = 3.0 MN, and (<b>c</b>) <span class="html-italic">V</span> = 5.0 MN.</p>
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<p>The bending moment along depth under different vertical pressures: (<b>a</b>) <span class="html-italic">V</span> = 1.0 MN, (<b>b</b>) <span class="html-italic">V</span> = 3.0 MN, and (<b>c</b>) <span class="html-italic">V</span> = 5 MN.</p>
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<p>The vertical load–settlement curves of a pile at mudline under different horizontal loads.</p>
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<p>The horizontal displacement along with depth under different horizontal loads: (<b>a</b>) <span class="html-italic">H</span> = 0.1 MN, (<b>b</b>) <span class="html-italic">H</span> = 0.3 MN, and (<b>c</b>) <span class="html-italic">H</span> = 0.5 MN.</p>
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<p>The (<b>a</b>) uplift load–vertical displacement and (<b>b</b>) uplift load–horizontal displacement curves of a pile at mudline at different lateral loads.</p>
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<p>Bearing capacity envelopes of piles under <span class="html-italic">V</span>–<span class="html-italic">H</span> loads.</p>
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22 pages, 6736 KiB  
Article
Enhanced Anti-Rollover Control for Commercial Vehicles Under Dynamic Lateral Interferences
by Jin Rong, Tong Wu, Junnian Wang, Jing Peng, Xiaojun Yang, Yang Meng and Liang Chu
Designs 2024, 8(6), 121; https://doi.org/10.3390/designs8060121 - 15 Nov 2024
Viewed by 351
Abstract
Commercial vehicles frequently experience lateral interferences, such as crosswinds or side slopes, during extreme maneuvers like emergency steering and high-speed driving due to their high centroid. These interferences reduce vehicle stability and increase the risk of rollover. Therefore, this study takes a bus [...] Read more.
Commercial vehicles frequently experience lateral interferences, such as crosswinds or side slopes, during extreme maneuvers like emergency steering and high-speed driving due to their high centroid. These interferences reduce vehicle stability and increase the risk of rollover. Therefore, this study takes a bus as the carrier and designs an anti-rollover control strategy based on mixed-sensitivity and robust H controller. Specifically, a 7-DOF vehicle dynamics model is introduced, and the factors influencing vehicle rollover are analyzed. Based on this, to minimize excessive intervention in the vehicle’s dynamic characteristics, the lateral velocity, roll angle, and roll rate are recorded at the vehicle’s rollover threshold as desired values. The lateral load transfer rate (LTR) is chosen as the evaluation index, and the required additional yaw moment is determined and distributed to the wheels for anti-rollover control. Furthermore, to verify the effectiveness of the proposed anti-rollover control strategy, a co-simulation platform based on MATLAB/Simulink and TruckSim is developed. Various dynamic lateral interferences (side winds with different changing trends and wind speeds) are introduced, and the fishhook and J-turn maneuvers are selected to analyze and compare the proposed control strategy with a fuzzy logic algorithm. The results indicate that the maximum LTR of the vehicle is reduced by 0.11. Additionally, the lateral acceleration and yaw rate in the steady state are reduced by more than 1.8 m/s² and 15°, respectively, enhancing the vehicle’s lateral stability. Full article
(This article belongs to the Topic Vehicle Dynamics and Control, 2nd Edition)
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<p>7-DOF vehicle dynamics model: (<b>a</b>) the lateral motion and yaw of the vehicle body and (<b>b</b>) the roll motion of the vehicle body.</p>
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<p>Vehicle dynamics comparison validation: (<b>a</b>) steering angle, (<b>b</b>) slip angle, (<b>c</b>) yaw rate, and (<b>d</b>) roll angle.</p>
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<p>Simulation curves under different wheel angles: (<b>a</b>) lateral velocity, (<b>b</b>) yaw rate, (<b>c</b>) roll angle, and (<b>d</b>) roll rate.</p>
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<p>Simulation curves at different vehicle speeds: (<b>a</b>) lateral velocity, (<b>b</b>) yaw rate, (<b>c</b>) roll angle, and (<b>d</b>) roll rate.</p>
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<p>Structure of anti-rollover control strategy. Where, <math display="inline"><semantics> <mrow> <msup> <mrow> <msub> <mi>ω</mi> <mrow> <mi>r</mi> <mi>d</mi> </mrow> </msub> </mrow> <mo>∗</mo> </msup> </mrow> </semantics></math>, <math display="inline"><semantics> <mrow> <msup> <mrow> <msub> <mi>v</mi> <mi>y</mi> </msub> </mrow> <mo>∗</mo> </msup> </mrow> </semantics></math>, <math display="inline"><semantics> <mrow> <msup> <mi>φ</mi> <mo>∗</mo> </msup> </mrow> </semantics></math> and <math display="inline"><semantics> <mrow> <msup> <mover accent="true"> <mi>φ</mi> <mo>˙</mo> </mover> <mo>∗</mo> </msup> </mrow> </semantics></math> are the desired value of yaw rate, lateral velocity, roll angle and roll rate, respectively.</p>
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<p>Generalized system.</p>
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<p>Relationship between wheel braking force and vehicle yaw moment.</p>
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<p>Relationship between cylinder pressure and wheel braking torque.</p>
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<p>Co-simulation platform.</p>
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<p>The membership function of inputs and output: (<b>a</b>) input and (<b>b</b>) output.</p>
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<p>Simulation results of fishhook maneuver: (<b>a</b>) steering wheel angle, (<b>b</b>) lateral velocity, (<b>c</b>) yaw rate, (<b>d</b>) roll angle, (<b>e</b>) roll rate, and (<b>f</b>) vehicle velocity.</p>
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<p>Simulation results of fishhook maneuver considering lateral interference: (<b>a</b>) wind speed, (<b>b</b>) lateral velocity, (<b>c</b>) yaw rate, (<b>d</b>) roll angle, (<b>e</b>) roll rate, and (<b>f</b>) vehicle velocity.</p>
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<p>Simulation diagram of J-turn working maneuver: (<b>a</b>) steering wheel angle, (<b>b</b>) lateral velocity, (<b>c</b>) yaw rate, (<b>d</b>) roll angle, (<b>e</b>) roll rate, and (<b>f</b>) vehicle velocity.</p>
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<p>Simulation diagram of lateral interference intervention under J-turn maneuver: (<b>a</b>) wind speed, (<b>b</b>) lateral velocity, (<b>c</b>) yaw rate, (<b>d</b>) roll angle, (<b>e</b>) roll rate, and (<b>f</b>) vehicle velocity.</p>
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34 pages, 9113 KiB  
Article
Investigation of Debonding Effect in Internal Replacement Pipe System Under Lateral Loading
by Tri C. M. Tien, Allan Manalo, Patrick Dixon, Warna Karunasena, Hamid Ahmadi, Shanika Kiriella, Ahmad Salah and Brad P. Wham
Appl. Sci. 2024, 14(22), 10509; https://doi.org/10.3390/app142210509 - 14 Nov 2024
Viewed by 600
Abstract
The aging pipeline infrastructure around the world necessitates immediate rehabilitation. Internal replacement pipe (IRP) is a trenchless system offering a versatile and cost-effective solution across a variety of industries, including oil, natural gas, water, and wastewater. As a structural pipeline repair system, IRPs [...] Read more.
The aging pipeline infrastructure around the world necessitates immediate rehabilitation. Internal replacement pipe (IRP) is a trenchless system offering a versatile and cost-effective solution across a variety of industries, including oil, natural gas, water, and wastewater. As a structural pipeline repair system, IRPs are subject to lateral deformation because of surface traffic loading. The present study evaluates the impact of adhesion between the host pipe and the IRP, with a focus on assessing the debonding effect on the behavior of the repair system under lateral deformation and bending. This was achieved using a comprehensive approach, including experimental, numerical, and analytical techniques. Varying levels of adhesive strength resulting from different methods of surface preparation were considered. The effectiveness of the IRP system on both discontinuous host pipes with various crack widths and continuous host pipes was also investigated. The results demonstrate that adhesive strength exerts a significant influence on the repair system, especially in the case of narrow circumferential cracks, while its impact on the continuous system is minimal. For optimal performance, it is essential to choose adhesives that possess sufficient shear strength while also accounting for the required debonding length. This approach ensures that minor discontinuities are effectively controlled, thereby enhancing the system′s fatigue life. The reliable determination of the maximum allowable shear strength for the adhesive or the debonding length can ensure that it does not negatively affect fatigue life. The findings presented in this study offer new insights into the development of trenchless repair techniques that can enhance system performance and extend service life. Full article
(This article belongs to the Section Mechanical Engineering)
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Figure 1
<p>A demonstration of the deflection test pipe system: (<b>a</b>) schematic and (<b>b</b>) physical setup. Courtesy of images from the University of Colorado Boulder [<a href="#B17-applsci-14-10509" class="html-bibr">17</a>].</p>
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<p>A demonstration of the deflection test pipe system: (<b>a</b>) schematic and (<b>b</b>) physical setup. Courtesy of images from the University of Colorado Boulder [<a href="#B17-applsci-14-10509" class="html-bibr">17</a>].</p>
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<p>The four-point bending FEA modeling and its components for the single pipe deflection validation.</p>
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<p>Meshing configurations for the single pipe deflection validation.</p>
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<p>The modeling of the IRP system: (<b>a</b>) whole view and (<b>b</b>) discontinuity view.</p>
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<p>Stress and strain curve for ALTRA10 IRP.</p>
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<p>Meshing configurations for IRP system deflection.</p>
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<p>Validation of load-displacement relationship in the 12.7 mm (0.5 inch) discontinuous IRP system.</p>
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<p>Shear stress results on the adhesive layer of the 12.7 mm (0.5 inch) discontinuous IRP system at a maximum load of 69.1 kN (15.5 kip) (magnified 18 times).</p>
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<p>Pull-off stress results on the adhesive layer of the 12.7 mm (0.5 inch) discontinuous IRP system at a maximum load of 69.1 kN (15.5 kip) (magnified 18 times).</p>
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<p>Equivalent stress results on the IRP of the 12.7 mm (0.5 inch) discontinuous IRP system at the system at 69.1 kN (15.5 kip) (magnified 25 times).</p>
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<p>FEA modeling and constraints for the lap-shear test in accordance with ASTM D5868.</p>
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<p>Meshing configuration of the lap-shear testing.</p>
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<p>The debonding progression of the adhesive layer with load increment.</p>
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<p>The force-displacement relationship of the load head during the adhesive validation test, as determined through numerical simulations.</p>
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<p>Equivalent stress results of a 12.7 mm (0.5 inch) discontinuous IRP system employing the debonding mechanism at a maximum load of 69.1 kN (15.5 kip) (magnified 29 times).</p>
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<p>Equivalent stress distribution in the IRP of a 12.7 mm (0.5 inch) discontinuous IRP system employing the debonding mechanism at a maximum load of 69.1 kN (15.5 kip) (magnified 13 times).</p>
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<p>Progressive debonding in the adhesive layer at various load levels for a 12.7 mm (0.5 inch) discontinuous IRP system (true scale).</p>
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<p>Progressive debonding in the adhesive layer at various load levels for a 12.7 mm (0.5 inch) discontinuous IRP system (true scale).</p>
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<p>Validation of the load-displacement relationship in the 12.7 mm (0.5 inch) discontinuous IRP system incorporating the debonding mechanism.</p>
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<p>Lateral loading versus deflection of the 0.5 in discontinuous steel IRP system under different levels of adhesive strengths.</p>
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<p>Average debonded length of the adhesive layer at the maximum load of 71.2 kN (16 kip) for 12.7 mm (0.5 inch) discontinuous IRP system versus different levels of adhesive shear strength.</p>
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<p>Normal stress results in the longitudinal direction of the IRP layer of a 12.7 mm (0.5 inch) discontinuous IRP system employing the adhesive shear strength of 20 MPa (2.9 ksi) at a fatigue load of 21.4 kN (4.8 kip) (magnified 74 times).</p>
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<p>Normal stress distribution in the longitudinal direction of the bottom line of the 12.7 mm (0.5 inch) discontinuous IRP system utilizing the debonding mechanism under a fatigue load of 21.4 kN (4.8 kip).</p>
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<p>Normal stress versus number of cycles to failure of ALTRA10 IRP [<a href="#B26-applsci-14-10509" class="html-bibr">26</a>].</p>
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<p>Photos of IRP specimen ′before′ and ′after′ mechanical aging tests: (<b>a</b>) internal and (<b>b</b>) cut apart joint. Courtesy images from Cornell University [<a href="#B27-applsci-14-10509" class="html-bibr">27</a>].</p>
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<p>Photos of IRP specimen ′before′ and ′after′ mechanical aging tests: (<b>a</b>) internal and (<b>b</b>) cut apart joint. Courtesy images from Cornell University [<a href="#B27-applsci-14-10509" class="html-bibr">27</a>].</p>
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<p>Equivalent stress results of the IRP layer of a 12.7 mm (0.5 inch) discontinuous IRP system employing the adhesive shear strength of 20 MPa (2.9 ksi) at a fatigue load of 71.2 kN (16 kip) (magnified 13 times).</p>
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<p>Equivalent stress distribution in the longitudinal direction of the bottom line of the 12.7 mm (0.5 inch) discontinuous IRP system utilizing the debonding mechanism under a maximum load of 71.2 kN (16 kip).</p>
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<p>Equivalent stress results of a 152.4 mm (6 inch) discontinuous IRP system employing the adhesive shear strength of 40 MPa (5.8 ksi) at a load of 56.9 kN (12.8 kip) (magnified 7.2 times).</p>
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<p>Lateral loading versus deflection of the 152.4 mm (6 inch) discontinuous IRP system under different levels of adhesive strengths.</p>
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<p>Average debonding length of the adhesive layer at the maximum load of 56.9 kN (12.8 kip) for 152.4 mm (6 inch) discontinuous IRP system versus different levels of adhesive shear strength.</p>
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<p>Normal stress results in the longitudinal direction of the IRP layer of a 152.4 mm (6 inch) discontinuous IRP system employing the adhesive shear strength of 20 MPa (2.9 ksi) at a fatigue load of 21.4 kN (4.8 kip) (magnified 23 times).</p>
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<p>Normal stress distribution in the longitudinal direction of the bottom line of the 12.7 mm (0.5 inch) discontinuous IRP system utilizing the debonding mechanism under a fatigue load of 21.4 kN (4.8 kip).</p>
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<p>Equivalent stress results in the IRP layer of a 152.4 mm (6 inch) discontinuous IRP system employing the adhesive shear strength of 20 MPa (2.9 ksi) at a maximum load of 56.9 kN (12.8 kip) (magnified 5.6 times).</p>
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<p>Equivalent stress results in the IRP layer of a 152.4 mm (6 inch) discontinuous IRP system employing the adhesive shear strength of 5 MPa (0.7 ksi) at a maximum load of 56.9 kN (12.8 kip) (magnified 4.5 times).</p>
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<p>Equivalent stress results in the IRP layer of a 152.4 mm (6 inch) discontinuous IRP system employing the adhesive shear strength of 40 MPa (5.8 ksi) at a maximum load of 56.9 kN (12.8 kip) (magnified 6.6 times).</p>
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<p>Local deformation failure in the IRP layer of a 152.4 mm (6 inch) discontinuous IRP system at a maximum load of 54.7 kN (12.3 kip). Courtesy data from the University of Colorado Boulder.</p>
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<p>Lateral loading versus deflection of the three IRP systems: continuous, 12.7 mm (0.5 inch), and 152.4 mm (6 inch) discontinuous.</p>
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<p>Shear stress results in the adhesive layer of a continuous IRP system employing the adhesive shear strength of 20 MPa (2.9 ksi) at a maximum load of 71.2 kN (magnified 98 times).</p>
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<p>Normal stress results in the IRP layer of a continuous IRP system employing the adhesive shear strength of 20 MPa (2.9 ksi) at a fatigue load of 71.2 kN (magnified 270 times).</p>
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<p>Equivalent stress results in the IRP layer of a continuous IRP system employing the adhesive shear strength of 20 MPa (2.9 ksi) at a maximum load of 71.2 kN (16 kip) (magnified 98 times).</p>
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18 pages, 3894 KiB  
Article
The Effect of a Single Temporomandibular Joint Soft Tissue Therapy on Cervical Spine Mobility, Temporomandibular Joint Mobility, Foot Load Distribution, and Body Balance in Women with Myofascial Pain in the Temporomandibular Joint Area—A Randomized Controlled Trial
by Iwona Sulowska-Daszyk, Paulina Handzlik-Waszkiewicz and Sara Gamrot
Appl. Sci. 2024, 14(22), 10397; https://doi.org/10.3390/app142210397 - 12 Nov 2024
Viewed by 522
Abstract
In contemporary times, a significant portion of the population experiences symptoms of temporomandibular joint (TMJ) dysfunction. The objective of this study was to evaluate the effects of a single-session TMJ soft tissue therapy on the TMJ and cervical spine mobility as well as [...] Read more.
In contemporary times, a significant portion of the population experiences symptoms of temporomandibular joint (TMJ) dysfunction. The objective of this study was to evaluate the effects of a single-session TMJ soft tissue therapy on the TMJ and cervical spine mobility as well as on body balance and the foot load distribution. This study was a parallel-group, randomized, controlled trial with a 1:1 allocation ratio. Fifty women aged 20–30 years diagnosed with myofascial pain in the TMJ area were included in the study and divided into two groups. The experimental group received TMJ soft tissue therapy. The following research tools were used: a Hogetex electronic caliper, a CROM Deluxe, and a FreeMed Base pedobarographic platform. In the experimental group, an increase in mobility within all assessed jaw and cervical spine movements was observed. This change was statistically significant (p < 0.05) for lateral movement to the left, abduction, and protrusion of the jaw (an increase of 10.32%, 7.07%, and 20.92%, respectively) and for extension, lateral bending to the right and left, and rotation to the right and left, of the cervical spine (an increase of 7.05%, 7.89%, 10.44%, 4.65%, and 6.55%, respectively). In the control group, no significant differences were observed. No significant changes were observed in the load distribution and body balance assessment. A single session of TMJ soft tissue therapy increases jaw and cervical spine mobility but does not impact body balance or foot load distribution in static conditions in women diagnosed with myofascial pain in the TMJ area. Full article
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<p>CONSORT flow diagram.</p>
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<p>The Hogetex electronic caliper.</p>
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<p>Cervical Range-of-Motion instrument.</p>
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<p>Posturographic examination in bipedal standing position.</p>
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<p>Posturographic examination in single-leg standing position.</p>
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<p>Temporalis muscle relaxation: trigger point therapy, external technique.</p>
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<p>Masseter muscle relaxation: myofascial release, external technique.</p>
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<p>Pterygoid muscle relaxation: internal technique.</p>
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<p>Masseter muscle relaxation: trigger point release in the attachment area, internal technique.</p>
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11 pages, 278 KiB  
Article
One-Side Weight Sports and the Impact of Their Load on the Feet and the Occurrence of Posture Disorders in Professional Football and Handball Players
by Matúš Kozel, Gabriela Škrečková, Marina Potašová, Peter Kutiš and Lenka Ondrušková
Appl. Sci. 2024, 14(22), 10387; https://doi.org/10.3390/app142210387 - 12 Nov 2024
Viewed by 431
Abstract
The aim of this study was to evaluate center of pressure (CoP) changes in unilateral sports and examine how these changes affect the athlete’s feet, ankles, knees or posture. The study sample consisted of 40 professional male players (age: 19.4 ± 2.08 years; [...] Read more.
The aim of this study was to evaluate center of pressure (CoP) changes in unilateral sports and examine how these changes affect the athlete’s feet, ankles, knees or posture. The study sample consisted of 40 professional male players (age: 19.4 ± 2.08 years; height: 165.78 ± 4.92 cm; weight: 59.04 ± 4.02 kg; BMI: 21.57 ± 2.22; foot size: 40.9 ± 1.6), divided by type of sport into group H—handball (n: 20) and group F—football (n: 20). To evaluate the monitored parameters, we used the instrumental diagnostic techniques: 3D laser footscan, baropodometric platform FreeStep and 2D Videography. We found no significant differences between the groups in the loading of the right and left foot (F: 8.3 ± 4.22; H: 7.7 ± 6.1) (p = 0.410). We found a significant difference in the load on the front and back of the left (p = 0.0079) and right foot (p = 0.0210) depending on the type of sport performed. Maximum and mean values of CoP (g/cm2) showed statistically significant differences depending on the sport performed (p < 0.0001). The shift in CoP (mm) from the norm depending on the sport performed was confirmed in the latero-lateral direction (p = 0.003), but not in the antero-posterior direction (p = 0.320). We found a difference in the angulation of the knees and heels depending on the sport played. Handball players showed higher knee varosity/valgosity (p = 0.015) and heel values than football players (p = 0.002). The handball players also confirmed a worse postural load and initial forward posture. The one-sided sports, handball and football, showed negative effects on the athlete’s movement system. These changes were more pronounced in handball players. Proper training programs should be applied in athletes’ daily routine to improve the negative effects of unilateral sports. Full article
(This article belongs to the Special Issue Advances in Foot Biomechanics and Gait Analysis)
18 pages, 5445 KiB  
Article
Monitoring of Non-Lame Horses and Horses with Unilateral Hindlimb Lameness at Rest with the Aid of Accelerometers
by Anja Uellendahl, Johannes P. Schramel, Alexander Tichy and Christian Peham
Sensors 2024, 24(22), 7203; https://doi.org/10.3390/s24227203 - 11 Nov 2024
Viewed by 395
Abstract
The aim of this study was to determine whether horses exhibiting unilateral hindlimb lameness unload (rest) the lame limb more than the contralateral limb. The resting/unloading of the hindlimbs and the time spent lying down were measured using accelerometers. Ten non-lame horses and [...] Read more.
The aim of this study was to determine whether horses exhibiting unilateral hindlimb lameness unload (rest) the lame limb more than the contralateral limb. The resting/unloading of the hindlimbs and the time spent lying down were measured using accelerometers. Ten non-lame horses and 20 lame horses were recruited for participation and monitored for 11 h overnight with accelerometers (MSR145, sampling rate: 1 Hz, and measuring range: ±15 g) attached to the lateral metatarsal and metacarpal regions of each limb. Metatarsal and metacarpal orientation were used to determine whether the limb was unloaded (rested) or loaded, respectively, or whether the horses were lying down. The relation of resting time between non-lame and lame limbs (non-lame/lame: 0.85 ± 1.2) of the lame horses differed significantly (p = 0.035) from that of the non-lame horses (right/left: 1.08 ± 0.47). Non-lame horses rested their hindlimbs evenly (left: 15 ± 10%; right: 17 ± 16%). Horses with unilateral hindlimb lameness unloaded the lame limb longer (lame limb: 61.8 ± 25.3%, non-lame limb: 38.2 ± 25.3%) than their contralateral limb. The lame horses (13 ± 11%) lay down longer (p = 0.012) than the non-lame horses (3 ± 6%). The degree of lameness determined by the participating veterinarians (Vet Score) (r = −0.691, p < 0.01) and the asymmetry evaluated by the lameness locator (ALL) (r = −0.426, p = 0.019) correlated with the resting ratio (rest time ratio). Both factors were also correlated with the time spent lying down (Vet Score (r = 0.364, p = 0.048) and the ALL (r = 0.398, p = 0.03)). The ALL and VET Score were significantly correlated (r = 0.557, p = 0.01). The results of this study provide a good baseline for future research into how individual resting patterns may help to detect pain. Full article
(This article belongs to the Section Physical Sensors)
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<p>The method of attaching the accelerometers (MSR145, sampling rate: 1 Hz, and measuring range: ±15 g) to the equine extremity laterally in the mid-region of the metacarpal and metatarsal bone with a Velcro fastener and one layer of elastic band.</p>
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<p>Comparison of the percentage distribution of the hindlimb resting time between limbs for lame and non-lame horses participating in this study to measure loading and unloading patterns of hindlimbs in lame horses compared to non-lame horses. Horse 6 is marked as an outlier in the first two boxplots (shown by the circles and number 6).</p>
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<p>Comparison and percentage distribution of the time spent lying down between the lame horses and the non-lame horses participating in this study to measure loading and unloading patterns in lame horses and non-lame horses. Horses 7 and 14 are marked as outliers in the right boxplot (shown by asterisks and the numbers 7 and 14).</p>
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<p>Objectively measured asymmetry between 20 horses with unilateral hindlimb lameness and 10 non-lame horses participating in this study to evaluate loading and unloading patterns in lame limbs compared to non-lame limbs (measured in millimeters (mm)).</p>
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<p>Illustration of the z-axis displacement when the position of a horse’s limb changes [<a href="#B8-sensors-24-07203" class="html-bibr">8</a>].</p>
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<p>Angle Calculation Basics [<a href="#B8-sensors-24-07203" class="html-bibr">8</a>].</p>
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<p>Calculating the angle to the vertical position [<a href="#B8-sensors-24-07203" class="html-bibr">8</a>].</p>
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<p>Coordinate axes in relation to a horse’s body.</p>
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<p>Coordinate transformation.</p>
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