Nothing Special   »   [go: up one dir, main page]

You seem to have javascript disabled. Please note that many of the page functionalities won't work as expected without javascript enabled.
 
 
Sign in to use this feature.

Years

Between: -

Subjects

remove_circle_outline
remove_circle_outline
remove_circle_outline
remove_circle_outline
remove_circle_outline
remove_circle_outline
remove_circle_outline
remove_circle_outline
remove_circle_outline

Journals

remove_circle_outline
remove_circle_outline
remove_circle_outline
remove_circle_outline
remove_circle_outline
remove_circle_outline
remove_circle_outline
remove_circle_outline
remove_circle_outline
remove_circle_outline
remove_circle_outline

Article Types

Countries / Regions

remove_circle_outline
remove_circle_outline
remove_circle_outline
remove_circle_outline
remove_circle_outline
remove_circle_outline
remove_circle_outline
remove_circle_outline

Search Results (1,487)

Search Parameters:
Keywords = vertical displacement

Order results
Result details
Results per page
Select all
Export citation of selected articles as:
21 pages, 5507 KiB  
Article
Load-Bearing Performance of Precast Piles with Integrated Side Drainage Channels in Coastal Soft Soil
by Shu-Hao Hu, Yue-Bao Deng, Shan Yu and Ri-Hong Zhang
Sustainability 2025, 17(5), 2324; https://doi.org/10.3390/su17052324 - 6 Mar 2025
Viewed by 178
Abstract
To accelerate the dissipation of excess pore water pressure, enhance the bearing capacity of piles, and mitigate long-term settlement in soft ground, a novel green and lowcarbon pile foundation technology, termed the precast drainage pile (PDP) technology, is proposed. This innovative approach integrated [...] Read more.
To accelerate the dissipation of excess pore water pressure, enhance the bearing capacity of piles, and mitigate long-term settlement in soft ground, a novel green and lowcarbon pile foundation technology, termed the precast drainage pile (PDP) technology, is proposed. This innovative approach integrated precast pipe piles with prefabricated vertical drains (PVDs) attached to their sides. The piles were installed using static pile pressing and were subsequently subjected to vacuum-induced negative pressure to facilitate soil consolidation, which enhances the resource utilization rate of pile foundations and promotes the sustainable utilization of soft soil foundations. To investigate the bearing characteristics of the PDP, this study combined the shear displacement method for piles with the consolidation theory of soft soil foundations. A calculation model for the load-settlement behavior of precast piles, accounting for the influence of vacuum-induced soil consolidation, was derived, establishing a method for analyzing the load transfer mechanism of PDPs. The reliability of the theoretical model was validated through comparisons with engineering test results. Building on this foundation, the influence of factors such as consolidation period and pile length on the bearing characteristics of PDPs was analyzed. The results demonstrated that, compared to a 10 m precast pile without drainage, the ultimate bearing capacity of single piles with drainage durations of 3, 7, 14, and 28 days increased by 7.3%, 12.7%, 20.3%, and 29.6%, respectively. Furthermore, under a 7-day drainage condition, the bearing capacity of piles with lengths of 10 m, 20 m, and 30 m increased by 12.7%, 12.8%, and 13.1%, respectively. Overall, the findings of this study provide a theoretical basis for the research, development, and design calculations of this new sustainable pile technology. Full article
Show Figures

Figure 1

Figure 1
<p>Precast drainage pile (PDP): (<b>a</b>) on-site construction of PDP; (<b>b</b>) schematic diagram of PDP.</p>
Full article ">Figure 2
<p>Consolidation model for single PVD surrounding the precast pile.</p>
Full article ">Figure 3
<p>Diagram of deformation between pile and the surrounding soil.</p>
Full article ">Figure 4
<p>Force analysis diagram for pile element.</p>
Full article ">Figure 5
<p>Load transfer function for pile side.</p>
Full article ">Figure 6
<p>Load transfer function for pile tip.</p>
Full article ">Figure 7
<p>Schematic diagram of elastic and plastic zones for pile–soil interaction.</p>
Full article ">Figure 8
<p>Calculation flow of load–settlement curve.</p>
Full article ">Figure 9
<p>Schematic diagram of field test scheme for case one.</p>
Full article ">Figure 10
<p>Static load tests for case one.</p>
Full article ">Figure 11
<p>Comparison of Q-S curves between measurement and theoretical value for case one.</p>
Full article ">Figure 12
<p>Field construction for case two.</p>
Full article ">Figure 13
<p>Comparison of Q-S curves between measurement and theoretical value for case two.</p>
Full article ">Figure 14
<p>Variation of load–settlement curves under different drainage periods.</p>
Full article ">Figure 15
<p>Load-bearing capacity and increase proportion with drainage periods.</p>
Full article ">Figure 16
<p>Comparisons of load–settlement curves under different pile lengths.</p>
Full article ">Figure 17
<p>Load-bearing capacity and increased proportion with pile lengths.</p>
Full article ">
28 pages, 9825 KiB  
Article
Study on the Application and Deformation Characteristics of Construction Waste Recycled Materials in Highway Subgrade Engineering
by Yuan Mei, Hongping Lu, Xueyan Wang, Bingyu Zhou, Ziyang Liu and Lu Wang
Buildings 2025, 15(5), 835; https://doi.org/10.3390/buildings15050835 - 6 Mar 2025
Viewed by 87
Abstract
It is difficult to meet environmental requirements via the coarse treatment methods of landfilling and open-air storage of construction waste. At the same time, the consumption of building materials in highway engineering is enormous. Using construction waste as a filling material for proposed [...] Read more.
It is difficult to meet environmental requirements via the coarse treatment methods of landfilling and open-air storage of construction waste. At the same time, the consumption of building materials in highway engineering is enormous. Using construction waste as a filling material for proposed roads has become a research hotspot in recent years. This paper starts with basic performance tests of recycled construction waste materials, and then moves on to laboratory experiments conducted to obtain the road performance of the recycled materials, the testing of key indicators of post-construction filling quality of the roadbed, and analyses of the deformation pattern of roadbed filled with construction waste. Additionally, the ABAQUS finite element software was used to establish a numerical model for roadbed deformation and analyze the roadbed deformation under different compaction levels and vehicle load conditions. The experimental results show that the recycled material has a moisture content of 8.5%, water absorption of 11.73%, and an apparent density of 2.61 g/cm3, while the liquid limit of fine aggregates is 20% and the plasticity index is 5.4. Although the physical properties are slightly inferior to natural aggregates, its bearing ratio (25–55%) and low expansion characteristics meet the requirements for high-grade highway roadbed filling materials. The roadbed layer with a loose compaction of 250 mm, after eight passes of rolling, showed a settlement difference of less than 5 mm, with the loose compaction coefficient stabilizing between 1.15 and 1.20. Finite element simulations indicated that the total settlement of the roadbed stabilizes at 20–30 mm, and increasing the compaction level to 96% can reduce the settlement by 2–4%. Vehicle overload causes a positive correlation between the vertical displacement and shear stress in the base layer, suggesting the need to strengthen vehicle load control. The findings provide theoretical and technical support for the large-scale application of recycled construction waste materials in roadbed engineering. Full article
(This article belongs to the Topic Sustainable Building Materials)
Show Figures

Figure 1

Figure 1
<p>Technical route diagram.</p>
Full article ">Figure 2
<p>Boundary moisture content test.</p>
Full article ">Figure 3
<p>Particle gradation curve.</p>
Full article ">Figure 4
<p>Standardized compaction test procedure.</p>
Full article ">Figure 5
<p>Effect of moisture content change on compacted specimens. (<b>a</b>) Water content 10%. (<b>b</b>) Water content 12%. (<b>c</b>) Water content 14%. (<b>d</b>) Water content 16%. (<b>e</b>) Water content 18%.</p>
Full article ">Figure 6
<p>Relationship between dry density and moisture content of coarse and fine aggregates in different proportions.</p>
Full article ">Figure 7
<p>Relationship between mixture content and compaction test results.</p>
Full article ">Figure 8
<p>California Bearing Ratio test procedure. (<b>a</b>) Specimen preparation. (<b>b</b>) Immersion of the specimen in water. (<b>c</b>) Specimen under pressure. (<b>d</b>) Specimen destruction.</p>
Full article ">Figure 9
<p>Test diagram of construction waste recycled materials for subgrade fill. (<b>a</b>) Test section. (<b>b</b>) Design of highway subgrade sections.</p>
Full article ">Figure 10
<p>Construction process of construction waste subgrade.</p>
Full article ">Figure 11
<p>Layout of compaction test points. (<b>a</b>) Compaction test cross-section. (<b>b</b>) Compaction test plan. (<b>c</b>) Compaction test site layout.</p>
Full article ">Figure 12
<p>EVD values corresponding to different compaction levels.</p>
Full article ">Figure 13
<p>Beckman beam measurement point layout and detection. (<b>a</b>) Elevation view of the detection point. (<b>b</b>) Plan view of detection points.</p>
Full article ">Figure 14
<p>Observation point layout.</p>
Full article ">Figure 15
<p>Settlement variation curve of the subgrade during the construction period due to construction waste. (<b>a</b>) Settlement variation in the left side of the subgrade. (<b>b</b>) Settlement variation in the right side of the subgrade. (<b>c</b>) Settlement of the subgrade cross-section.</p>
Full article ">Figure 16
<p>Numerical model of the subgrade.</p>
Full article ">Figure 17
<p>Boundary conditions.</p>
Full article ">Figure 18
<p>Displacement cloud map of roadbed. (<b>a</b>) Horizontal displacement of roadbed (X direction). (<b>b</b>) Vertical displacement of roadbed (Y direction).</p>
Full article ">Figure 19
<p>Roadbed deformation cloud map corresponding to different compaction degrees. (<b>a</b>) Cloud map of roadbed settlement when compaction degree is 90%. (<b>b</b>) Cloud map of roadbed settlement when compaction degree is 93%. (<b>c</b>) Cloud map of roadbed settlement when compaction degree is 96%.</p>
Full article ">Figure 20
<p>Loading area.</p>
Full article ">Figure 21
<p>Grid division.</p>
Full article ">Figure 22
<p>Vertical displacement corresponding to different loads. (<b>a</b>) Vertical displacement corresponding to a load of 100 kN. (<b>b</b>) Vertical displacement corresponding to a load of 120 kN. (<b>c</b>) Vertical displacement corresponding to a load of 160 kN. (<b>d</b>) Vertical displacement corresponding to a load of 200 kN.</p>
Full article ">Figure 23
<p>Load and vertical displacement of roadbed. (<b>a</b>) Vertical displacement curve corresponding to different loads on the top surface of the base. (<b>b</b>) Vertical displacement curve corresponding to different loads at the top surface of the base layer (0.38 m). (<b>c</b>) Vertical displacement curve corresponding to different loads at the top of the roadbed (0.58 m).</p>
Full article ">Figure 24
<p>Effect of load on shear stress.</p>
Full article ">
27 pages, 15905 KiB  
Article
Tracking the Seismic Deformation of Himalayan Glaciers Using Synthetic Aperture Radar Interferometry
by Sandeep Kumar Mondal, Rishikesh Bharti and Kristy F. Tiampo
Remote Sens. 2025, 17(5), 911; https://doi.org/10.3390/rs17050911 - 5 Mar 2025
Viewed by 106
Abstract
The Himalayan belt, formed due to the Cenozoic convergence between the Eurasian and Indian craton, acts as a storehouse of large amounts of strain, resulting in large earthquakes from the Western to the Eastern Himalayas. Glaciers also occur over a major portion of [...] Read more.
The Himalayan belt, formed due to the Cenozoic convergence between the Eurasian and Indian craton, acts as a storehouse of large amounts of strain, resulting in large earthquakes from the Western to the Eastern Himalayas. Glaciers also occur over a major portion of the high-altitude Himalayan region. The impact of earthquakes can be easily studied in the plains and plateaus with the help of well-distributed seismogram networks and these regions’ accessibility is helpful for field- and lab-based studies. However, earthquakes triggered close to high-altitude Himalayan glaciers are tough to investigate for the impact over glaciers and glacial deposits. In this study, we attempt to understand the impact of earthquakes on and around Himalayan glaciers in terms of vertical displacement and coherence change using space-borne synthetic aperture radar (SAR). Eight earthquake events of various magnitudes and hypocenter depths occurring in the vicinity of Himalayan glacial bodies were studied using C-band Sentinel1-A/B SAR data. Differential interferometric SAR (DInSAR) analysis is applied to capture deformation of the glacial surface potentially related to earthquake occurrence. Glacial displacement varies from −38.9 mm to −5.4 mm for the 2020 Tibet earthquake (Mw 5.7) and the 2021 Nepal earthquake (Mw 4.1). However, small glacial and ground patches processed separately for vertical displacements reveal that the glacial mass shows much greater seismic displacement than the ground surface. This indicates the possibility of the presence of potential site-specific seismicity amplification properties within glacial bodies. A reduction in co-seismic coherence around the glaciers is observed in some cases, indicative of possible changes in the glacial moraine deposits and/or vegetation cover. The effect of two different seismic events (the 2020 and 2021 Nepal earthquakes) with different hypocenter depths but with the same magnitude at almost equal distances from the glaciers is assessed; a shallow earthquake is observed to result in a larger impact on glacial bodies in terms of vertical displacement. Earthquakes may induce glacial hazards such as glacial surging, ice avalanches, and the failure of moraine-/ice-dammed glacial lakes. This research may be able to play a possible role in identifying areas at risk and provide valuable insights for the planning and implementation of measures for disaster risk reduction. Full article
(This article belongs to the Section Environmental Remote Sensing)
Show Figures

Figure 1

Figure 1
<p>Study area map showing (<b>A</b>) world map with geographical boundary of India, tectonic plate boundaries, and (<b>B</b>) the Himalayan glaciers, tectonic lineaments, Indian–Eurasian plate boundaries, and selected earthquake epicenters (sources: Esri, HERE, Garmin International, and others as mentioned on the map).</p>
Full article ">Figure 2
<p>Process diagram for estimating atmospherically corrected vertical displacement and coherence of ground and glaciers.</p>
Full article ">Figure 3
<p>Map for 2020 Tibet earthquake (M<sub>w</sub> 5.7) showing (<b>A</b>) orthorectified phase interferogram, (<b>B</b>) vertical displacement, (<b>C</b>) vertical displacement for coherence ≥ 0.6, co-seismic image, (<b>D</b>) orthorectified phase interferogram at epicenter location, (<b>E</b>) vertical displacement at epicenter location, and (<b>F</b>) vertical displacement for coherence ≥ 0.6, co-seismic image.</p>
Full article ">Figure 4
<p>Map showing orthorectified layer of (<b>A</b>) difference between pre- and co-seismic coherence, (<b>B</b>,<b>C</b>) vertical displacement (mm) within glacial bodies derived from unwrapped phase interferogram for 2020 Tibet earthquake (M<sub>w</sub> 5.7) within black and brown rectangular regions, respectively.</p>
Full article ">Figure 5
<p>Map showing (<b>A</b>) orthorectified vertical displacement and (<b>B</b>) vertical displacement with co-seismic coherence ≥ 0.6 for 2020 Leh earthquake (M<sub>w</sub> 5.3).</p>
Full article ">Figure 6
<p>Map showing (<b>A</b>) difference between pre- and co-seismic coherence and (<b>B</b>) vertical displacement (mm) within glacial bodies derived from unwrapped phase interferogram for 2020 Leh earthquake (M<sub>w</sub> 5.3).</p>
Full article ">Figure 7
<p>Map showing (<b>A</b>) glacial and ground subset regions, (<b>B</b>) vertical displacement (mm) within glacial subset, and (<b>C</b>) vertical displacement (mm) within ground subset derived from unwrapped phase interferogram for 2020 Leh earthquake (M<sub>w</sub> 5.3).</p>
Full article ">Figure 8
<p>Map showing (<b>A</b>) orthorectified vertical displacement and (<b>B</b>) vertical displacement with co-seismic coherence ≥ 0.6 for 2017 Thang earthquake (M<sub>w</sub> 5.2).</p>
Full article ">Figure 9
<p>Map showing (<b>A</b>) difference between pre- and co-seismic coherence and (<b>B</b>) vertical displacement (mm) within glacial bodies derived from unwrapped phase interferogram for 2017 Thang earthquake (M<sub>w</sub> 5.2).</p>
Full article ">Figure 10
<p>Map showing (<b>A</b>) glacial and ground subset regions, (<b>B</b>) vertical displacement (mm) within glacial subset-I, (<b>C</b>) vertical displacement (mm) within glacial subset-II, (<b>D</b>) vertical displacement (mm) within ground subset-I, and (<b>E</b>) vertical displacement (mm) within ground subset-II.</p>
Full article ">Figure 11
<p>Map showing (<b>A</b>) orthorectified vertical displacement and (<b>B</b>) vertical displacement with co-seismic coherence ≥ 0.6 for 2021 Joshimath earthquake (M<sub>w</sub> 4.5).</p>
Full article ">Figure 12
<p>Map showing (<b>A</b>,<b>B</b>,<b>D</b>) difference between pre- and co-seismic coherence within the influence area, black and red rectangular regions, respectively, and (<b>C</b>,<b>E</b>) vertical displacement (mm) within glacial bodies derived from unwrapped phase interferogram for 2021 Joshimath earthquake (M<sub>w</sub> 4.5) for black and red rectangular regions, respectively.</p>
Full article ">Figure 13
<p>Map showing (<b>A</b>) glacial and ground subset regions, (<b>B</b>) vertical displacement (mm) within glacial subset, and (<b>C</b>) vertical displacement (mm) within ground subset derived from unwrapped phase interferogram for 2021 Joshimath earthquake (M<sub>w</sub> 4.5).</p>
Full article ">Figure 14
<p>Relation between (<b>A</b>) <span class="html-italic">IR</span> and M<sub>w</sub> of earthquakes triggered at hypocenter depth of 10 km and relation between (<b>B</b>) <span class="html-italic">IR<sub>N</sub></span> and M<sub>w</sub> of earthquakes triggered close to hypocenter depth of 49.8 km.</p>
Full article ">Figure 15
<p>Map showing (<b>A</b>) difference between pre- and co-seismic coherence and (<b>B</b>) vertical displacement (mm) within influence area derived from unwrapped phase interferogram for 2017 Sikkim earthquake (M<sub>w</sub> 4.2).</p>
Full article ">Figure 16
<p>Map showing difference between pre- and co-seismic coherence for 2018 Sikkim earthquake (M<sub>w</sub> 4.4).</p>
Full article ">Figure 17
<p>Relation between IRN and M<sub>w</sub> of earthquakes triggered close to hypocenter depth of 35 km.</p>
Full article ">Figure 18
<p>Map showing (<b>A</b>) difference between pre- and co-seismic coherence and (<b>B</b>) vertical displacement (mm) within an influence area derived from unwrapped phase interferogram for 2020 Nepal earthquake (M<sub>w</sub> 4.1).</p>
Full article ">Figure 19
<p>Map showing (<b>A</b>) difference between pre- and co-seismic coherence and (<b>B</b>–<b>D</b>) vertical displacement (mm) within influence area derived from unwrapped phase interferogram for 2021 Nepal earthquake (M<sub>w</sub> 4.1).</p>
Full article ">Figure 20
<p>Map showing vertical displacement (mm) within glacial bodies of common study zone for (<b>A</b>) 2020 Nepal earthquake (M<sub>w</sub> 4.1) and (<b>B</b>) 2021 Nepal earthquake (M<sub>w</sub> 4.1).</p>
Full article ">Figure 21
<p>False-color-composite (FCC) map of the common study zone in Nepal.</p>
Full article ">
29 pages, 8539 KiB  
Article
Three-Dimensional FEM Analysis of the Protective Effects of Isolation Piles on Tunnels Under Adjacent Excavations
by Libo Xu, Junneng Ye, Yanming Yao, Chi Liu and Xiaoli Liu
Appl. Sci. 2025, 15(5), 2673; https://doi.org/10.3390/app15052673 - 2 Mar 2025
Viewed by 291
Abstract
Isolation piles are critical for mitigating excavation-induced tunnel displacements, yet two unresolved challenges persist in tunnel engineering: (1) controversies regarding the influence of key parameters (e.g., pile head depth, pile length, and pile-to-pit distance) on their performance, and (2) insufficient understanding of the [...] Read more.
Isolation piles are critical for mitigating excavation-induced tunnel displacements, yet two unresolved challenges persist in tunnel engineering: (1) controversies regarding the influence of key parameters (e.g., pile head depth, pile length, and pile-to-pit distance) on their performance, and (2) insufficient understanding of the effects on both horizontal and vertical displacement control of tunnel. These challenges stem from the current research focus on isolated displacement components or simplified scenarios, which fails to address the complex interactions between key parameters and the deformation mechanisms. To address these gaps, this study proposes a hybrid validation framework integrating a three-dimensional finite element model (HS-Small constitutive model) with field monitoring data. A concept of “control efficiency” is introduced to quantify the effectiveness of isolation piles, complemented by a parametric sensitivity analysis framework. By synergizing the mirror image method and statistical theory, the research reveals a dual-path control mechanism involving displacement blocking and tunnel geometric reconfiguration. The findings advance the state of the art by resolving controversies over critical parameters and establishing a unified theoretical framework for coupled displacement control, providing actionable insights for optimizing isolation pile design in engineering practice. Full article
(This article belongs to the Special Issue New Challenges in Urban Underground Engineering)
Show Figures

Figure 1

Figure 1
<p>A workflow diagram of this study.</p>
Full article ">Figure 2
<p>Stress–strain relationship in a standard drained triaxial test.</p>
Full article ">Figure 3
<p>Schematic layout of the project.</p>
Full article ">Figure 4
<p>Profile of the relative positions of the foundation pit, isolation pile, and tunnel.</p>
Full article ">Figure 5
<p>Engineering geological cross-section.</p>
Full article ">Figure 6
<p>Three-dimensional finite element model: (<b>a</b>) model, (<b>b</b>) structure, and (<b>c</b>) mesh map.</p>
Full article ">Figure 7
<p>Evolutions of the horizontal displacement of the retaining wall (C1) and the depth.</p>
Full article ">Figure 8
<p>Evolutions of the horizontal displacement of the soil at a distance of 15 m away from the excavation (C2) and the depth.</p>
Full article ">Figure 9
<p>Evolutions of the horizontal displacement (<b>a</b>) and vertical displacement (<b>b</b>) of the tunnel with the monitoring point along the tunnel.</p>
Full article ">Figure 10
<p>Comparison between the predicted and measured values.</p>
Full article ">Figure 11
<p>Simplified finite-element model for parametric study.</p>
Full article ">Figure 12
<p>Three-dimensional finite-element model for parametric analysis.</p>
Full article ">Figure 13
<p>Relationships between control efficiency and pile position (x = 2 m) for varying pile lengths: 5 m (<b>a</b>), 10 m (<b>b</b>), 15 m (<b>c</b>), 20 m (<b>d</b>), 25 m (<b>e</b>), 30 m (<b>f</b>), and 35 m (<b>g</b>).</p>
Full article ">Figure 14
<p>Relationships between control efficiency and pile position (x = 5.5 m) for varying pile lengths: 5 m (<b>a</b>), 10 m (<b>b</b>), 15 m (<b>c</b>), 20 m (<b>d</b>), 25 m (<b>e</b>), 30 m (<b>f</b>), and 35 m (<b>g</b>).</p>
Full article ">Figure 15
<p>Relationships between control efficiency and pile position (x = 9 m) for varying pile lengths: 5 m (<b>a</b>), 10 m (<b>b</b>), 15 m (<b>c</b>), 20 m (<b>d</b>), 25 m (<b>e</b>), 30 m (<b>f</b>), and 35 m (<b>g</b>).</p>
Full article ">Figure 16
<p>Evolutions of the maximum horizontal displacement of soil at the isolation piles with the depth of the pile head for cases with x = 2 (<b>a</b>) and 9 m (<b>b</b>).</p>
Full article ">Figure 17
<p>Evolutions of vertical (<b>a</b>) and horizontal (<b>b</b>) displacement of the tunnel with the center angel (α).</p>
Full article ">Figure 18
<p>Schematic diagram of the tunnel deformation due to excavation with and without isolation piles.</p>
Full article ">Figure 19
<p>Relationships between control efficiency and pile head depth for cases with x = 2 (<b>a</b>), 5.5 (<b>b</b>), and 9 m (<b>c</b>).</p>
Full article ">Figure 20
<p>Relationships between control efficiency and pile length for cases with x = 2 (<b>a</b>), 5.5 (<b>b</b>), and 9 m (<b>c</b>).</p>
Full article ">Figure 21
<p>Relationships between the maximum horizontal displacement of soil at isolation piles and the position of the pile head and toe for cases with x = 2 m.</p>
Full article ">Figure 22
<p>Schematic diagram of the tunnel deformation due to excavation with various pile lengths.</p>
Full article ">Figure 23
<p>Relationships between vertical control efficiency and distance between the foundation pit and isolation piles for varying pile lengths: 10 m (<b>a</b>), 15 m (<b>b</b>), 20 m (<b>c</b>), 25 m (<b>d</b>), 30 m (<b>e</b>), 35 m (<b>f</b>), and 40 m (<b>g</b>).</p>
Full article ">Figure 24
<p>Relationships between horizontal control efficiency and distance between the foundation pit and isolation piles for varying pile lengths: 10 m (<b>a</b>), 15 m (<b>b</b>), 20 m (<b>c</b>), 25 m (<b>d</b>), 30 m (<b>e</b>), 35 m (<b>f</b>), and 40 m (<b>g</b>).</p>
Full article ">Figure 25
<p>Variation pattern of the control efficiency of isolation piles for a constant pile length (<b>a</b>) and pile toe depth (<b>b</b>).</p>
Full article ">
16 pages, 4497 KiB  
Article
Experimental Investigation on the Application of Polymer Agents in Offshore Sandstone Reservoirs: Optimization Design for Enhanced Oil Recovery
by Yanyue Li, Changlong Liu, Yaqian Zhang, Baoqing Xue, Jinlong Lv, Chuanhui Miao, Yiqiang Li and Zheyu Liu
Polymers 2025, 17(5), 673; https://doi.org/10.3390/polym17050673 - 2 Mar 2025
Viewed by 265
Abstract
The conventional polymer gel has high initial viscosity and short gelation time, making it difficult to meet the requirements of deep profile control in offshore reservoirs with large well spacing and strong heterogeneity. This paper evaluates the performance and core plugging capacity of [...] Read more.
The conventional polymer gel has high initial viscosity and short gelation time, making it difficult to meet the requirements of deep profile control in offshore reservoirs with large well spacing and strong heterogeneity. This paper evaluates the performance and core plugging capacity of novel functional polymer gels and microspheres to determine the applicability of core permeability ranges. On the heterogeneous core designed based on the reservoir characteristics of Block B oilfield, optimization was conducted separately for the formulation, dosage, and slug combinations of the polymer gel/microsphere. Finally, oil displacement experiments using polymer and microsphere combinations were conducted on vertically and planar heterogeneous cores to simulate reservoir development effects. The experimental results show the novel functional polymer gel exhibits slow gelation with high gel strength, with viscosity rapidly increasing four days after aging, ultimately reaching a gel strength of 74,500 mPa·s. The novel functional polymer gel and polymer microsphere can effectively plug cores with permeabilities below 6000 mD and 2000 mD, respectively. For heterogeneous cores with an average permeability of 1000 mD, the optimal polymer microsphere has a concentration of 4000 mg/L and a slug size of 0.3 PV; for heterogeneous cores with an average permeability of 4000 mD, the optimal functional polymer gel has a concentration of 7500 mg/L and a slug size of 0.1 PV. In simulations of vertically and planarly heterogeneous reservoirs, the application of polymer agent increases the oil recovery factor by 53% and 38.7% compared to water flooding. This realizes the gradual and full utilization of layers with high, medium, and low permeability. Full article
(This article belongs to the Special Issue New Studies of Polymer Surfaces and Interfaces: 2nd Edition)
Show Figures

Figure 1

Figure 1
<p>Preparation and Gelation Mechanism of Functional polymer gel.</p>
Full article ">Figure 2
<p>Physical and design drawings of cores.</p>
Full article ">Figure 3
<p>Viscosity curve of polymer agent.</p>
Full article ">Figure 4
<p>Production dynamic curves of polymer microsphere at different concentrations.</p>
Full article ">Figure 5
<p>Production dynamic curves of functional polymer gel at different concentrations.</p>
Full article ">Figure 6
<p>Production dynamic curves under different slug sizes of polymer microsphere.</p>
Full article ">Figure 7
<p>Production dynamic curves under different slug sizes of functional polymer gel.</p>
Full article ">Figure 8
<p>Production dynamic curves in heterogeneous reservoirs.</p>
Full article ">Figure 9
<p>Production dynamic curves in planar heterogeneous reservoirs.</p>
Full article ">Figure 9 Cont.
<p>Production dynamic curves in planar heterogeneous reservoirs.</p>
Full article ">Figure 10
<p>Planar distribution map of pressure. (<span class="html-fig-inline" id="polymers-17-00673-i001"><img alt="Polymers 17 00673 i001" src="/polymers/polymers-17-00673/article_deploy/html/images/polymers-17-00673-i001.png"/></span> Injection Well; <span class="html-fig-inline" id="polymers-17-00673-i002"><img alt="Polymers 17 00673 i002" src="/polymers/polymers-17-00673/article_deploy/html/images/polymers-17-00673-i002.png"/></span> Production Well).</p>
Full article ">
16 pages, 4377 KiB  
Article
Analysis of the Impact of the New Two-Lane Shield Tunnel Underpass on the Existing Tunnels
by Jinkui Li, Xinxia Fang and Yu Yang
Appl. Sci. 2025, 15(5), 2642; https://doi.org/10.3390/app15052642 - 28 Feb 2025
Viewed by 300
Abstract
To address the issue of vertical settlement in existing tunnels beneath newly constructed two-lane tunnels, and to mitigate further impacts on tunnel operations, it is essential to investigate the effect of tunnel construction on the integrity of the existing tunnel structure. A calculation [...] Read more.
To address the issue of vertical settlement in existing tunnels beneath newly constructed two-lane tunnels, and to mitigate further impacts on tunnel operations, it is essential to investigate the effect of tunnel construction on the integrity of the existing tunnel structure. A calculation formula for the vertical displacement of the existing tunnel is derived by simplifying the calculation model and employing a two-stage analysis method. A three-dimensional numerical model of the double-line shield tunnel beneath the existing tunnel of Dalian Metro Line 4 is established using Midas GTS NX finite element software 2021(v1. 1). The study focuses on the influence of the new tunnel’s excavation on the existing tunnel, examining how various parameters in the shield construction process affect the settlement. Through comparative analysis of theoretical calculations, numerical simulations, and engineering monitoring data, the results indicate that the calculated displacement settlement trends align closely with the numerical simulation and are consistent with the field monitoring data. The findings provide valuable insights for the development of effective protection measures for existing tunnels during shield tunnel construction. Full article
Show Figures

Figure 1

Figure 1
<p>Schematic diagram of the calculation model.</p>
Full article ">Figure 2
<p>Schematic diagram of ground settlement underneath two-lane shield tunnel.</p>
Full article ">Figure 3
<p>Plane diagram of new Line 4 and existing Line 5.</p>
Full article ">Figure 4
<p>Geologic profile of underpass interval.</p>
Full article ">Figure 5
<p>3D numerical model.</p>
Full article ">Figure 6
<p>Cloud image of sedimentation deformation.</p>
Full article ">Figure 7
<p>Settlement of existing tunnels.</p>
Full article ">Figure 8
<p>Vertical deformation rule of left line monitoring point.</p>
Full article ">Figure 9
<p>Influence curve of tunneling pressure on existing tunnel deformation.</p>
Full article ">Figure 10
<p>Influence curve of grouting pressure on deformation of existing tunnels.</p>
Full article ">Figure 11
<p>Layout of monitoring sites. (<b>a</b>) Schematic layout of surface measurement points. (<b>b</b>) Schematic diagram of measuring points in the main section of the tunnel.</p>
Full article ">Figure 12
<p>Comparison of theoretical calculation, numerical simulation, and field data.</p>
Full article ">
18 pages, 1792 KiB  
Article
Similarity Index Values in Fuzzy Logic and the Support Vector Machine Method Applied to the Identification of Changes in Movement Patterns During Biceps-Curl Weight-Lifting Exercise
by André B. Peres, Tiago A. F. Almeida, Danilo A. Massini, Anderson G. Macedo, Mário C. Espada, Ricardo A. M. Robalo, Rafael Oliveira, João P. Brito and Dalton M. Pessôa Filho
J. Funct. Morphol. Kinesiol. 2025, 10(1), 84; https://doi.org/10.3390/jfmk10010084 - 28 Feb 2025
Viewed by 157
Abstract
Background/Objectives: Correct supervision during the performance of resistance exercises is imperative to the correct execution of these exercises. This study presents a proposal for the use of Morisita–Horn similarity indices in modelling with machine learning methods to identify changes in positional sequence [...] Read more.
Background/Objectives: Correct supervision during the performance of resistance exercises is imperative to the correct execution of these exercises. This study presents a proposal for the use of Morisita–Horn similarity indices in modelling with machine learning methods to identify changes in positional sequence patterns during the biceps-curl weight-lifting exercise with a barbell. The models used are based on the fuzzy logic (FL) and support vector machine (SVM) methods. Methods: Ten male volunteers (age: 26 ± 4.9 years, height: 177 ± 8.0 cm, body weight: 86 ± 16 kg) performed a standing barbell bicep curl with additional weights. A smartphone was used to record their movements in the sagittal plane, providing information about joint positions and changes in the sequential position of the bar during each lifting attempt. Maximum absolute deviations of movement amplitudes were calculated for each execution. Results: A variance analysis revealed significant deviations (p < 0.002) in vertical displacement between the standard execution and execution with a load of 50% of the subject’s body weight. Experts with over thirty years of experience in resistance-exercise evaluation evaluated the exercises, and their results showed an agreement of over 70% with the results of the ANOVA. The similarity indices, absolute deviations, and expert evaluations were used for modelling in both the FL system and the SVM. The root mean square error and R-squared results for the FL system (R2 = 0.92, r = 0.96) were superior to those of the SVM (R2 = 0.81, r = 0.79). Conclusions: The use of FL in modelling emerges as a promising approach with which to support the assessment of movement patterns. Its applications range from automated detection of errors in exercise execution to enhancing motor performance in athletes. Full article
(This article belongs to the Special Issue Biomechanical Analysis in Physical Activity and Sports)
Show Figures

Figure 1

Figure 1
<p>Scheme for capturing videos. Image partly generated by AI in: <a href="https://firefly.adobe.com/generate/images" target="_blank">https://firefly.adobe.com/generate/images</a> (accessed on 14 September 2024).</p>
Full article ">Figure 2
<p>Scatterplot of MH index values vs. absolute deviation values.</p>
Full article ">Figure 3
<p>Fuzzy Inference System model.</p>
Full article ">Figure 4
<p>Variation in the values of the absolute deviations at the system input.</p>
Full article ">Figure 5
<p>Comparison between original Morisita–Horn data and data obtained from the fuzzy logic model.</p>
Full article ">Figure 6
<p>Comparison between original Morisita–Horn data and data obtained by the coarse Gaussian SVM model.</p>
Full article ">
19 pages, 7509 KiB  
Article
Effects of Vertical Irregularity on Transverse Reinforcement Spacing in Reinforced Concrete Columns to Avoid Shear Failure Subjected to Seismic Behavior
by Hak-Jong Chang, Jae-Hyun Cho, Mun-Gi Kim and Jun-Hee Kim
Buildings 2025, 15(5), 785; https://doi.org/10.3390/buildings15050785 - 27 Feb 2025
Viewed by 172
Abstract
As a result of the 2017 Pohang earthquake, numerous piloti-type structures incurred damage, and the cause was attributed to the wide spacing of transverse reinforcement. Improper spacing of transverse reinforcement can lead to brittle failure of columns, potentially causing the collapse of buildings. [...] Read more.
As a result of the 2017 Pohang earthquake, numerous piloti-type structures incurred damage, and the cause was attributed to the wide spacing of transverse reinforcement. Improper spacing of transverse reinforcement can lead to brittle failure of columns, potentially causing the collapse of buildings. This study aimed to analyze the failure mode of columns where load and displacement are concentrated due to vertical irregularity, and to quantify the spacing of shear reinforcement according to the degree of vertical irregularity to prevent shear failure of the column. First, a vertically irregular frame with vertical irregularity and an RC moment frame with the same upper and lower structural systems was modeled, and the failure mode of the column was analyzed. In this paper, the failure modes were classified into shear failure, flexure–shear failure, and flexural failure based on the shear capacity ratio. The analysis results showed that in the case of vertical irregularity, the shear demand of the column was evaluated as high due to the high flexural stiffness of the horizontal members, and the failure mode of the column was classified as shear failure. The impact of the spacing of shear reinforcement on the shear strength of the structure was also examined. Next, an analysis was performed according to the degree of vertical irregularity by adjusting the thickness of the first-floor shear wall, and as a result, the proportion of the entire columns classified as shear failure increased as the vertical irregularity increased. It was confirmed that the minimum spacing of shear reinforcement of 150 mm specified in Korean standards becomes inadequate when the degree of vertical irregularity exceeds 2.6. At a vertical irregularity of 8.3, the spacing required to prevent shear failure decreased to 136 mm, which is 9.33% less than the minimum specified by the Korean standards. This indicates that even if the code’s minimum spacing is adhered to, shear failure can still occur in columns. In order to prevent shear failure of the column, the spacing of the shear reinforcement should be designed smaller, because the shear force increases as the vertical irregularity increases. For piloti-type structures with high horizontal irregularity, there is a need to design shear reinforcement narrower than the minimum standard to prevent shear failure of the column. Full article
(This article belongs to the Section Building Structures)
Show Figures

Figure 1

Figure 1
<p>Damage of the piloti-type structures in Korea due to the 2017 Pohang earthquake: (<b>a</b>) Damage on the 1st story; (<b>b</b>) Damage of the columns on the 1st story (picture taken by Structural System and Earthquake Engineering Laboratory (SSEEL) at Yonsei University).</p>
Full article ">Figure 2
<p>Concept of failure mode of the column decided by shear demand and shear strength.</p>
Full article ">Figure 3
<p>Shear strength by different parameters causation.</p>
Full article ">Figure 4
<p>Shear demand of columns depending on lateral stiffness ratio between beam and column.</p>
Full article ">Figure 5
<p>Model of vertically irregular frame and RC moment frame.</p>
Full article ">Figure 6
<p>Column’s shear DCR for VIF and MF with transverse rebar spacing of 200 mm.</p>
Full article ">Figure 7
<p>Shear demand by lateral stiffness due to vertical irregularity.</p>
Full article ">Figure 8
<p>Shear strength by vertical irregularity and transverse reinforcement spacing: (<b>a</b>) <math display="inline"><semantics> <mrow> <msub> <mrow> <mi>V</mi> </mrow> <mrow> <mi>c</mi> </mrow> </msub> </mrow> </semantics></math> of VIF and MF; (<b>b</b>) <math display="inline"><semantics> <mrow> <msub> <mrow> <mi>V</mi> </mrow> <mrow> <mi>s</mi> </mrow> </msub> </mrow> </semantics></math> varies on transverse reinforcement spacing; (<b>c</b>) ratios of <math display="inline"><semantics> <mrow> <msub> <mrow> <mi>V</mi> </mrow> <mrow> <mi>c</mi> </mrow> </msub> </mrow> </semantics></math>/<math display="inline"><semantics> <mrow> <msub> <mrow> <mi>V</mi> </mrow> <mrow> <mi>n</mi> </mrow> </msub> </mrow> </semantics></math> differ by transverse reinforcement spacing.</p>
Full article ">Figure 9
<p>Maximum base shear ratio of VIF and MF.</p>
Full article ">Figure 10
<p>Vertically irregular frame with shear wall on 1st floor: (<b>a</b>) floor plan; (<b>b</b>) elevation.</p>
Full article ">Figure 11
<p>Difference of vertical irregularity due to thickness of shear wall.</p>
Full article ">Figure 12
<p>Shear demand capacity ratio (DCR) according to vertical irregularity: (<b>a</b>) average shear DCR, (<b>b</b>) DCR distribution of each column according to vertical irregularity.</p>
Full article ">Figure 13
<p>Appropriate transverse reinforcement spacing for preventing shear failure in columns.</p>
Full article ">Figure 14
<p>Analysis of actual Piloti-type structures: (<b>a</b>) vertical irregularity of structures; (<b>b</b>) transverse reinforcement spacing for bending failure.</p>
Full article ">
33 pages, 13351 KiB  
Article
Modeling and Investigation of Long-Term Performance of High-Rise Pile Cap Structures Under Scour and Corrosion
by Shilei Niu, Zhongxiang Liu, Tong Guo, Anxin Guo and Sudong Xu
J. Mar. Sci. Eng. 2025, 13(3), 450; https://doi.org/10.3390/jmse13030450 - 26 Feb 2025
Viewed by 262
Abstract
High-rise pile cap structures, such as sea-crossing bridges, suffer from long-term degradation due to continuous corrosion and scour, which seriously endangers structural safety. However, there is a lack of research on this topic. This study focused on the long-term performance and dynamic response [...] Read more.
High-rise pile cap structures, such as sea-crossing bridges, suffer from long-term degradation due to continuous corrosion and scour, which seriously endangers structural safety. However, there is a lack of research on this topic. This study focused on the long-term performance and dynamic response of bridge pile foundations, considering scour and corrosion effects. A refined modeling method for bridge pile foundations, considering scour-induced damage and corrosion-induced degradation, was developed by adjusting nonlinear soil springs and material properties. Furthermore, hydrodynamic characteristics and long-term performance, including hydrodynamic phenomena, wave force, energy, displacement, stress, and acceleration responses, were investigated through fluid–structure coupling analysis and pile–soil interactions. The results show that the horizontal wave forces acting on the high-rise pile cap are greater than the vertical wave forces, with the most severe wave-induced damage occurring in the wave splash zone. Steel and concrete degradation in the wave splash zone typically occurs sooner than in the atmospheric zone. The total energy of the structure at each moment under load is equal to the sum of internal energy and kinetic energy. Increased corrosion time and scour depth result in increased displacement and stress at the pile cap connection. The long-term dynamic response is mainly influenced by the second-order frequency (62 Hz). Full article
(This article belongs to the Special Issue Wave Loads on Offshore Structure)
Show Figures

Figure 1

Figure 1
<p>Corrosion damage to the substructures of sea-crossing bridges.</p>
Full article ">Figure 2
<p>Wave-cap numerical flume model.</p>
Full article ">Figure 3
<p>Detailed drawing of the grid division.</p>
Full article ">Figure 4
<p>Detailed dimensional drawing of the pier cap (unit: cm).</p>
Full article ">Figure 5
<p>Time variation curve of the wave surface at the wave height monitoring location.</p>
Full article ">Figure 6
<p>Detailed structural diagrams for bridge and high-rise pile cap.</p>
Full article ">Figure 7
<p>Constitutive relationship of concrete and steel reinforcement in numerical modeling.</p>
Full article ">Figure 8
<p>Finite element model of high-rise pile cap.</p>
Full article ">Figure 9
<p><span class="html-italic">P</span>–<span class="html-italic">y</span> curves of soil layers at different depths.</p>
Full article ">Figure 10
<p>Simulation diagram of scouring process.</p>
Full article ">Figure 11
<p>Performance degradation curves of steel bars and concrete.</p>
Full article ">Figure 12
<p>Flow diagram of dynamic response analysis of pile cap structure.</p>
Full article ">Figure 13
<p>Wave generation process.</p>
Full article ">Figure 14
<p>Pressure cloud diagram.</p>
Full article ">Figure 15
<p>The distribution of turbulent kinetic energy on the fluid–structure interface at 32 s.</p>
Full article ">Figure 16
<p>Wave force–time history curve.</p>
Full article ">Figure 17
<p>Energy analysis of high-rise pile cap.</p>
Full article ">Figure 18
<p>The displacement–time history curve and the Fourier analysis of the cap used for 8 years.</p>
Full article ">Figure 19
<p>Transverse displacement cloud diagram of the cap under 6 m scouring.</p>
Full article ">Figure 20
<p>Displacement–time history curve of pile top and Fourier analysis.</p>
Full article ">Figure 21
<p>Maximum transverse displacement and fitting curve for pile top #1.</p>
Full article ">Figure 22
<p>Location of the concern points.</p>
Full article ">Figure 23
<p>Transverse stress–time relationship.</p>
Full article ">Figure 24
<p>Vertical stress distribution diagram of pile #1 at 5.8 s.</p>
Full article ">Figure 25
<p>Transverse acceleration–time relationship of the cap under different scour depths.</p>
Full article ">Figure 26
<p>Fourier spectrum of transverse acceleration.</p>
Full article ">Figure 27
<p>Scatter plot of the maximum transverse acceleration of the cap.</p>
Full article ">
15 pages, 5117 KiB  
Article
In Situ Study on Vertical Compressive Bearing Characteristics of Rooted Bored Piles
by Chao Yang, Guoliang Dai, Weiming Gong, Shuang Xi, Mingxing Zhu and Shaolei Huo
Buildings 2025, 15(5), 707; https://doi.org/10.3390/buildings15050707 - 23 Feb 2025
Viewed by 262
Abstract
In situ vertical load field tests were carried out on two bored piles used in the Chizhou Highway Bridge across the Yangtze River, both of which were rooted piles. Based on the test results, such as those on the relationship between the load [...] Read more.
In situ vertical load field tests were carried out on two bored piles used in the Chizhou Highway Bridge across the Yangtze River, both of which were rooted piles. Based on the test results, such as those on the relationship between the load and settlement, axial force distribution, and the relationship between shaft friction and pile–soil relative displacement, the vertical load transfer mechanics of the rooted piles were analyzed. The results showed that the load-carrying curves of the rooted piles vary gradually and also that the rooted piles exhibit the bearing characteristics of friction piles because the loads at the pile tips are less than 15% of the total bearing capacity of the piles. The slope of the axial force distribution curve of the rooted piles first increased at the upper interface and then decreased at the lower interface of the root-reinforced zone. The axial force of the rooted piles decreased faster in soil layers where the piles had roots, which can be explained by the fact that roots share the vertical load with piles and that roots improve the bearing properties of piles. Considering the shaft and end resistance of the roots on the piles, the relationship between load and settlement of the rooted piles was calculated by a three-line model based on the load transfer method. The results calculated from the model were in good agreement with the results from the tests. The results from the tests could inform the design and analysis of rooted piles. Full article
Show Figures

Figure 1

Figure 1
<p>Typical special-shaped piles. (<b>a</b>) Tapered pile. (<b>b</b>) Branch and plate pile. (<b>c</b>) Under-reamed piles. (<b>d</b>) Rooted piles.</p>
Full article ">Figure 2
<p>Layout of roots. (<b>a</b>) Roots on pile plane layout and physical objects. (<b>b</b>) Rebar of roots.</p>
Full article ">Figure 2 Cont.
<p>Layout of roots. (<b>a</b>) Roots on pile plane layout and physical objects. (<b>b</b>) Rebar of roots.</p>
Full article ">Figure 3
<p>Umbrella stacking load in field loading test.</p>
Full article ">Figure 4
<p>Vibrating wire steel strain gauge.</p>
Full article ">Figure 5
<p>Steel gauge layout of tested pile. (<b>a</b>) SZ-1. (<b>b</b>) SZ-2.</p>
Full article ">Figure 6
<p><span class="html-italic">Q</span>–<span class="html-italic">s</span> curves of test piles.</p>
Full article ">Figure 7
<p>Axial force distribution of test piles. (<b>a</b>) SZ-1 test pile. (<b>b</b>) SZ-2 test pile.</p>
Full article ">Figure 8
<p>Relationship between end resistance ratio and loading at pile top.</p>
Full article ">Figure 9
<p>Shaft resistance distribution curves of piles. (<b>a</b>) SZ-1 test pile. (<b>b</b>) SZ-2 test pile.</p>
Full article ">Figure 10
<p>Calculation model of load transfer method.</p>
Full article ">Figure 11
<p>Relationship between side resistance and pile–soil relative displacement.</p>
Full article ">Figure 12
<p>Sketch of <span class="html-italic">τ</span>–<span class="html-italic">s</span> curve fitted by three-line model.</p>
Full article ">Figure 13
<p>Comparison of calculated result and tested result. (<b>a</b>) <span class="html-italic">Q</span>–<span class="html-italic">s</span> curve. (<b>b</b>) Axial force distribution curve.</p>
Full article ">
22 pages, 8041 KiB  
Article
The Bearing Capacity Model of Pile Foundation with Hole-Drilling and Pile-Inserting Technology in Complex Geological Environments
by Yi Wang, Guoyun Lu, En Zhang, Cheng Zhao, Wei Wang and Fenghui Dong
Buildings 2025, 15(5), 703; https://doi.org/10.3390/buildings15050703 - 23 Feb 2025
Viewed by 313
Abstract
Karst geology creates a complex environment with diverse landforms, blurred boundaries, and multi-factor interactions. This paper presents a new drilling pile installation method: drill to a set depth, clean the hole, insert prefabricated piles, and drive or vibrate them to the target elevation. [...] Read more.
Karst geology creates a complex environment with diverse landforms, blurred boundaries, and multi-factor interactions. This paper presents a new drilling pile installation method: drill to a set depth, clean the hole, insert prefabricated piles, and drive or vibrate them to the target elevation. It suits tough geological conditions well. Pile foundations bear both axial and lateral eccentric loads. To explore prestressed high-strength concrete (PHC) pile foundations under eccentric vertical loads in karst areas, on-site bearing capacity tests were conducted. The results show that as load eccentricity increases, PHC pile foundation-bearing capacity drops notably. A finite element model was developed to analyze the stress and strain behavior of PHC pile foundations under eccentric loading in complex geological conditions, aiming to assess their bearing capacity and stability. Key findings include: (1) Under constant external load, the maximum displacement of the PHC pile foundation increases with greater load eccentricity. (2) Enhanced concrete strength reduces the maximum displacement of the pile foundation, while the peak stress remains stable. (3) The height of karst caves has a minimal impact on the bearing capacity and deformation of PHC pile foundations. These results highlight the importance of considering load eccentricity, concrete strength, and cave height in optimizing the design of PHC pile foundations for safety in complex geological settings. Full article
Show Figures

Figure 1

Figure 1
<p>Geological materials and geometric parameters around pile foundation.</p>
Full article ">Figure 2
<p>Pictures of the test site. (<b>a</b>) Process of implanting pile, (<b>b</b>) static loading process of pile foundation.</p>
Full article ">Figure 3
<p>(<b>a</b>) Static load Q-s curve, (<b>b</b>) maximum settlement and increasing rate vs. eccentricity ratio, and (<b>c</b>) bearing capacity and decreasing rate vs. eccentricity ratio of PHC500AB125 (the eccentricity ratio is defined as the ratio of the eccentricity distance of the external load to the length of the pile foundation).</p>
Full article ">Figure 4
<p>Finite element model.</p>
Full article ">Figure 5
<p>The settlement of pile foundation with eccentricity of 0.015: experimental test and finite element simulation.</p>
Full article ">Figure 6
<p>The maximum displacement of pile foundations under different eccentricities and different loading forces.</p>
Full article ">Figure 7
<p>The displacement contour of pile foundation under different eccentricities and different loading forces. (<b>a</b>) Eccentricity of 0.001, external load of 2000 kN; (<b>b</b>) eccentricity of 0.001, external load of 5000 kN; (<b>c</b>) eccentricity of 0.02, external load of 2000 kN; (<b>d</b>) eccentricity of 0.02, external load of 5000 kN.</p>
Full article ">Figure 8
<p>The maximum stress of pile foundations under different eccentricities and different loading forces.</p>
Full article ">Figure 9
<p>The stress contour of pile foundation under different eccentricities and different loading forces. (<b>a</b>) Eccentricity of 0.001, external load of 2000 kN. (<b>b</b>) Eccentricity of 0.001, external load of 5000 kN. (<b>c</b>) Eccentricity of 0.02, external load of 2000 kN. (<b>d</b>) Eccentricity of 0.02, external load of 5000 kN.</p>
Full article ">Figure 10
<p>The variation law of pile foundation displacement with eccentricity under design load.</p>
Full article ">Figure 11
<p>The variation law of pile foundation stress with eccentricity under design load.</p>
Full article ">Figure 12
<p>The maximum displacement of pile foundations under different elastic moduli and different loading forces (the eccentricity is 0.005).</p>
Full article ">Figure 13
<p>The maximum stress of pile foundations under different elastic moduli and different loading forces (the eccentricity is 0.005).</p>
Full article ">Figure 14
<p>The maximum displacement of pile foundations under different elastic moduli and different loading forces (the eccentricity is 0.01).</p>
Full article ">Figure 15
<p>The maximum stress of pile foundations under different elastic moduli and different loading forces (the eccentricity is 0.01).</p>
Full article ">Figure 16
<p>The displacement contour of pile foundations under different elastic moduli and different loading forces (the eccentricity is 0.01). (<b>a</b>) C30, external load of 2000 kN. (<b>b</b>) C30, external load of 5000 kN. (<b>c</b>) C55, external load of 2000 kN. (<b>d</b>) C55, external load of 5000 kN.</p>
Full article ">Figure 17
<p>The stress contour of pile foundations under different elastic moduli and different loading forces (the eccentricity is 0.01). (<b>a</b>) C30, external load of 2000 kN. (<b>b</b>) C30, external load of 5000 kN. (<b>c</b>) C55, external load of 2000 kN. (<b>d</b>) C55, external load of 5000 kN.</p>
Full article ">Figure 18
<p>The variation law of pile foundation displacement with concrete grade under design load.</p>
Full article ">Figure 19
<p>The variation law of pile foundation stresses with concrete grade under design load.</p>
Full article ">Figure 20
<p>The variation curve of the maximum displacement of pile foundation with the height of rock.</p>
Full article ">Figure 21
<p>The variation curve of the maximum stress of pile foundation with the height of rock.</p>
Full article ">Figure 22
<p>The augmentation rate of (<b>a</b>) maximum displacement and (<b>b</b>) maximum stress of pile foundations under different eccentricities and different loading forces.</p>
Full article ">Figure 23
<p>The reduction rate of (<b>a</b>) maximum displacement and (<b>b</b>) maximum stress of pile foundations under different concrete grades and different loading forces.</p>
Full article ">Figure 24
<p>The reduction rate of (<b>a</b>) maximum displacement and (<b>b</b>) maximum stress of pile foundations under different rock heights.</p>
Full article ">
17 pages, 9752 KiB  
Article
Hydroxyapatite Dental Inserts for Tooth Restoration: Stress and Displacement Analysis
by Maja Lezaja Zebic, Aleksandar Bodic, Djordje Veljovic, Tamara Matic, Jelena Carkic and Vladimir Milovanovic
J. Funct. Biomater. 2025, 16(3), 75; https://doi.org/10.3390/jfb16030075 - 20 Feb 2025
Viewed by 305
Abstract
Hydroxyapatite (HAP) inserts minimize restoration contraction by constituting a major part of the restoration; however, their effect on the relaxation of tooth tissues has not been previously tested. Finite element analysis was employed to estimate stress and displacement when HAP inserts with a [...] Read more.
Hydroxyapatite (HAP) inserts minimize restoration contraction by constituting a major part of the restoration; however, their effect on the relaxation of tooth tissues has not been previously tested. Finite element analysis was employed to estimate stress and displacement when HAP inserts with a thickness of 1.7 mm or 4.7 mm and a diameter of 4.7 mm were used to substitute for dentin. The volumetric contraction of the composite during polymerization, simulated through steady-state heat transfer analysis, yielded a contraction rate of 3.7%. Descriptive statistics revealed that the incorporation of HAP inserts reduced the displacement of dentin, enamel, and restoration caused by contraction by 44.4% to 66.7%, while maximal stress was reduced by 8.1% to 52%. Subsequent loading on the occlusal tooth surface showed that displacement values decreased by 12.1% to 33.3%, while maximum von Mises stress in enamel decreased by 32.8% to 40.6% with the use of HAP inserts. Although the maximum stress values in dentin were not significantly decreased (3% to 8.8%), the stress located at the bottom of the cavity was notably reduced, particularly in deep cavities at root canal entrances. The use of HAP inserts in restorative dentistry provides benefits for the preservation of prepared teeth, especially in preventing irreparable vertical root fractures of endodontically treated teeth. Full article
(This article belongs to the Special Issue Biomechanical Studies and Biomaterials in Dentistry)
Show Figures

Figure 1

Figure 1
<p>The right mandibular first molar was extracted from the jaw, along with adjacent bone, tested as a healthy tooth, and prepared for “Small” and “Big” cavity restorations. For the “C” groups, the restoration comprised only composite, while for the “I+C” groups, it included cement for insert cementation, the HAP insert itself, and the composite above the insert. A cross-section of the tested models, with marked components, is presented in the left portion of the figure.</p>
Full article ">Figure 2
<p>Boundary conditions.</p>
Full article ">Figure 3
<p>First and second loading phases were implemented in the simulation. The first phase involves simulating the composite polymerization contraction with the model to calculate the percentage of contraction. A tooth cross-section is presented, demonstrating the temperature used for contraction provocation. The second loading phase involves the contraction simulation followed by a load of 2000 N continuously distributed over the occlusal surface.</p>
Full article ">Figure 4
<p>Graphical presentations of stress detected in all model components for the control (Small C and Big C) and experimental (Small I+C and Big I+C) first loading phase, which simulated polymerization contraction alone are provided.</p>
Full article ">Figure 5
<p>Von Mises stress distribution in healthy teeth after load application and for other groups after multistep analysis is provided. The legend value bar on the left side is common for dentin and the restoration, while the legend value bar on the right side is specific to enamel.</p>
Full article ">Figure 6
<p>Maximal detected values of von Misses stress in dentin and enamel, for both loading phases with decrease percentages shown.</p>
Full article ">
14 pages, 7641 KiB  
Article
Accuracy Assessment of Ocean Tide Models in the Eastern China Marginal Seas Using Tide Gauge and GPS Data
by Junjie Wang and Xiufeng He
J. Mar. Sci. Eng. 2025, 13(3), 395; https://doi.org/10.3390/jmse13030395 - 20 Feb 2025
Viewed by 280
Abstract
Accurate ocean tide models are required to remove tidal loading effects in geophysical research. Beyond a mere intercomparison, the accuracy of eight modern global models (DTU10, EOT20, FES2014b, FES2022b, GOT4.10c, HAMTIDE11a, OSU12, TPXO10-atlas-v2) and one regional model (NAO99Jb) was assessed in the eastern [...] Read more.
Accurate ocean tide models are required to remove tidal loading effects in geophysical research. Beyond a mere intercomparison, the accuracy of eight modern global models (DTU10, EOT20, FES2014b, FES2022b, GOT4.10c, HAMTIDE11a, OSU12, TPXO10-atlas-v2) and one regional model (NAO99Jb) was assessed in the eastern China marginal seas (ECMSs) using geodetic measurements. This involved rigorous comparisons with the tidal constant measurements at 65 tide gauges and with the GPS-measured M2 vertical ocean tide loading (OTL) displacements at 22 sites. The selected models showed significant disagreements close to the coasts of eastern China and the western Korean Peninsula, where the largest discrepancy for the M2 constituent could exceed 30 cm. However, EOT20 and FES2014b provided relatively close results, differing by only about 15 cm in Hangzhou Bay. EOT20 compared more favourably than the others to the tidal constant measurements, with a root sum square (RSS) of 11.1 cm, and to the GPS-measured M2 vertical OTL displacements, with a root mean square (RMS) of 0.49 mm. In addition, to differentiate between ocean tide models with subtle discrepancies when comparing them with the OTL measurements, consideration of the asthenospheric anelasticity effect was necessary. Full article
(This article belongs to the Section Physical Oceanography)
Show Figures

Figure 1

Figure 1
<p>The M<sub>2</sub> cotidal map and bathymetry across the ECMSs based on the FES2014b ocean tide model and the ETOPO 1′ grid, respectively. The black lines represent the contours of equal amplitude with an interval of 30 cm. The white contour lines indicate the phase lag with a contour interval of 45°. The phase lag is zero on the dotted line, with the white arrow indicating the direction of increasing phase lag.</p>
Full article ">Figure 2
<p>The M<sub>2</sub> phasor differences between the nine selected ocean tide models.</p>
Full article ">Figure 3
<p>The M2 phasor differences between the measured and the modelled tidal constants for (<b>a</b>) DTU10, (<b>b</b>) EOT20, (<b>c</b>) FES2014b, (<b>d</b>) FES2022b, (<b>e</b>) GOT4.10c, (<b>f</b>) HAMTIDE11a, (<b>g</b>) NAO99Jb, (<b>h</b>) OSU12, (<b>i</b>) TPXO-atlas-v2. The circles and squares represent the tide gauges from the JODC and UHSLC and Zhang [<a href="#B39-jmse-13-00395" class="html-bibr">39</a>], respectively. Differences larger than 50 cm are labelled with their precise values.</p>
Full article ">Figure 4
<p>The RMS and RSS values (in centimetres) when comparing the tide gauge measurements and the individual ocean tide model.</p>
Full article ">Figure 5
<p>The phasor differences between the GPS-measured M<sub>2</sub> vertical OTL displacements and the predictions computed using the aPREM_M2 Green’s function for (<b>a</b>) DTU10, (<b>b</b>) EOT20, (<b>c</b>) FES2014b, (<b>d</b>) FES2022b, (<b>e</b>) GOT4.10c, (<b>f</b>) HAMTIDE11a, (<b>g</b>) NAO99Jb, (<b>h</b>) OSU12, (<b>i</b>) TPXO-atlas-v2. Differences larger than 2 mm are labelled with their precise values.</p>
Full article ">Figure 6
<p>The ranking of the ocean tide models according to the RMS agreements between the GPS-measured and the predicted M2 vertical OTL displacements using aPREM_M2 and PREM Green’s functions, respectively.</p>
Full article ">
19 pages, 4816 KiB  
Article
Thickness Model of the Adhesive on Spacecraft Structural Plate
by Yanhui Guo, Peibo Li, Yanpeng Chen, Xinfu Chi and Yize Sun
Aerospace 2025, 12(2), 159; https://doi.org/10.3390/aerospace12020159 - 19 Feb 2025
Viewed by 165
Abstract
This paper establishes a physical model for the non-contact rotary screen coating process based on a spacecraft structural plate and proposes a theoretical expression for the adhesive thickness of the non-contact rotary screen coating. The thickness of the adhesive is a critical factor [...] Read more.
This paper establishes a physical model for the non-contact rotary screen coating process based on a spacecraft structural plate and proposes a theoretical expression for the adhesive thickness of the non-contact rotary screen coating. The thickness of the adhesive is a critical factor influencing the quality of the optical solar reflector (OSR) adhesion. The thickness of the adhesive layer depends on the equivalent fluid height and the ratio of the fluid flow rate to the squeegee speed below the squeegee. When the screen and fluid remain constant, the fluid flow rate below the squeegee depends on the pressure at the tip of the squeegee. The pressure is also a function related to the deformation characteristics and speed of the squeegee. Based on the actual geometric shape of the wedge-shaped squeegee, the analytical expression for the vertical displacement of the squeegee is obtained as the actual boundary of the flow field. The analytical expression for the deformation angle of the squeegee is used to solve the contact length between the squeegee and the rotary screen. It reduces the calculation difficulty compared with the previous method. Based on the theory of rheology and fluid mechanics, the velocity distribution of the fluid under the squeegee and the expression of the dynamic pressure at the tip of the squeegee were obtained. The dynamic pressure at the tip of the squeegee is a key factor for the adhesive to pass through the rotary screen. According to the continuity equation of the fluid, the theoretical thickness expression of the non-contact rotary screen coating is obtained. The simulation and experimental results show that the variation trend of coating thickness with the influence of variables is consistent. Experimental and simulation errors compared to theoretical values are less than 5%, which proves the rationality of the theoretical expression of the non-contact rotary screen coating thickness under the condition of considering the actual squeegee deformation. The existence of differences proves that a small part of the colloid remains on the rotary screen during the colloid transfer process. The expression parameterizes the rotary screen coating model and provides a theoretical basis for the design of automatic coating equipment. Full article
(This article belongs to the Section Astronautics & Space Science)
Show Figures

Figure 1

Figure 1
<p>The actuator of the non-contact rotary screen coating process with measurement function: (<b>b</b>) shows the internal structure of the actuator in (<b>a</b>).</p>
Full article ">Figure 2
<p>The geometric model of the squeegee.</p>
Full article ">Figure 3
<p>The contours with FEM for deformed configurations of the squeegee. Initial and deformation states are presented by the solid border and color area, respectively.</p>
Full article ">Figure 4
<p>Deformation curve of the lower edge of the wedge part for the FEM and the analytical solution. (<b>a</b>) 40 N, 75<math display="inline"><semantics> <mrow> <mo>°</mo> </mrow> </semantics></math>; (<b>b</b>) 60 N, 75<math display="inline"><semantics> <mrow> <mo>°</mo> </mrow> </semantics></math>; (<b>c</b>) 40 N, 85<math display="inline"><semantics> <mrow> <mo>°</mo> </mrow> </semantics></math>; (<b>d</b>) 60 N, 85<math display="inline"><semantics> <mrow> <mo>°</mo> </mrow> </semantics></math>.</p>
Full article ">Figure 5
<p>Schematic of flow field division for the rotary screen printing process. <math display="inline"><semantics> <mrow> <msub> <mrow> <mi>X</mi> </mrow> <mrow> <mi>a</mi> <mi>q</mi> </mrow> </msub> </mrow> </semantics></math> is the accumulated length of adhesive in front of the squeegee. <math display="inline"><semantics> <mrow> <msub> <mrow> <mi>V</mi> </mrow> <mrow> <mi>p</mi> </mrow> </msub> </mrow> </semantics></math> is the adhesive flow speed under the squeegee. <math display="inline"><semantics> <mrow> <msub> <mrow> <mi>V</mi> </mrow> <mrow> <mi>s</mi> </mrow> </msub> </mrow> </semantics></math> is the printing speed of the squeegee and rotary screen. <math display="inline"><semantics> <mrow> <msub> <mrow> <mi>H</mi> </mrow> <mrow> <mi>d</mi> </mrow> </msub> </mrow> </semantics></math> is the adhesive layer thickness. <math display="inline"><semantics> <mrow> <msub> <mrow> <mi>H</mi> </mrow> <mrow> <mi>s</mi> </mrow> </msub> </mrow> </semantics></math> is the rotary screen thickness.</p>
Full article ">Figure 6
<p>Flow field model in front of the squeegee.</p>
Full article ">Figure 7
<p>Velocity distribution in the flow field: (<b>a</b>) Non-dimensional x-velocity of a vertical section under different non-dimensional pressure gradients; (<b>b</b>) velocity distribution of adhesive in the vertical section at x = X relative to the substrate.</p>
Full article ">Figure 8
<p>Partial schematic diagram (dashed square) and flow area of the rotary screen. a is the projection direction. The blue area is the mesh. the dashed square is the sampling area of the rotary screen.</p>
Full article ">Figure 9
<p>The dynamic pressure at the squeegee tip under different conditions. The pressure of the squeegee with (<b>a</b>) force, (<b>b</b>) angle, (<b>c</b>) elastic modulus, and (<b>d</b>) coating speed. The black line represents the given parameters, and the x-axis represents the changing parameters in the figure.</p>
Full article ">Figure 10
<p>Deposited thickness in relation to the thickness under the squeegee.</p>
Full article ">Figure 11
<p>Variations in thickness <math display="inline"><semantics> <mrow> <msub> <mrow> <mi>H</mi> </mrow> <mrow> <mi>d</mi> </mrow> </msub> </mrow> </semantics></math> under different parameters. (<b>a</b>) Variations in thickness with angle for different power-law indices. Variations in thickness with (<b>b</b>) force <span class="html-italic">F</span>, (<b>c</b>) screen thickness <math display="inline"><semantics> <mrow> <msub> <mrow> <mi>H</mi> </mrow> <mrow> <mi>s</mi> </mrow> </msub> </mrow> </semantics></math>, and (<b>d</b>) elastic modulus <span class="html-italic">E</span> at θ = 75°and 85° for different power-law indices <span class="html-italic">n</span>.</p>
Full article ">Figure 12
<p>Variation and fit equation in viscosity with shear rate.</p>
Full article ">Figure 13
<p>CFD simulation of rotary screen printing: (<b>a</b>) Boundary conditions of CFD model; (<b>b</b>) grid division of CFD model; volume fraction of the phase contours in (<b>c</b>) t = 0.2 and (<b>d</b>) t = 0.6 cases; (<b>e</b>) total pressure field; and (<b>f</b>) streamline pattern of velocity.</p>
Full article ">Figure 14
<p>Photo and scanning contours of rotary screen printing: (<b>a</b>) Coating test; (<b>b</b>) scanning image of the adhesive layer. Red is the thicker area, while blue is the thinner area.</p>
Full article ">Figure 15
<p>Comparison of the thickness <math display="inline"><semantics> <mrow> <msub> <mrow> <mi>H</mi> </mrow> <mrow> <mi>d</mi> </mrow> </msub> </mrow> </semantics></math> is obtained by the approximate analytical solution (red solid line), the experimental solution (black triangle), and the CFD solution (Blue Square). (<b>a</b>) Variation in thickness with force. (<b>b</b>) Variation in thickness with angle.</p>
Full article ">Figure 16
<p>Experimental and simulation errors compared to theoretical values.</p>
Full article ">
20 pages, 16736 KiB  
Article
Numerical Simulation of Mechanical Response of Tunnel Breakage in the Construction of Cross Passages by Mechanical Excavation Method Using Flat-Face Cutterhead
by Bingyi Li, Xianghong Li and Songyu Liu
Appl. Sci. 2025, 15(4), 2153; https://doi.org/10.3390/app15042153 - 18 Feb 2025
Viewed by 237
Abstract
Mechanical construction has gradually been applied in cross passages of metro lines, but more mechanical mechanisms should be revealed. The section between Jingrong Street Station and Kunjia Road Station in Suzhou Metro Line 11 adopts a mechanical construction method to construct a cross [...] Read more.
Mechanical construction has gradually been applied in cross passages of metro lines, but more mechanical mechanisms should be revealed. The section between Jingrong Street Station and Kunjia Road Station in Suzhou Metro Line 11 adopts a mechanical construction method to construct a cross passage. A novel flat-face cutterhead, which is different from curved cutter head is first used to cut and break the main tunnel in construction of cross passage. Based on the background of practical engineering, the finite element method was applied to simulate the breaking process of the main tunnel to explore the dynamic variation in the mechanical response of the segments cut by the flat-face cutterhead. The results indicate that the maximum vertical displacement caused by cutting mainly concentrates on the top of the fully cut rings. The maximum horizontal displacement occurs at the waist on the side of the tunnel portal in the semi-cut rings. The axial force level inside both types of segment rings reaches its peak after the tunnel is formed. The maximum axial force exists at the bottom and top of the fully cut ring and semi-cut ring, respectively. The change in the displacement around the portal is not substantial before the third stage, and it begins to increase significantly from the moment the concrete at the portal is penetrated. The existence of the pre-support system effectively controls the displacement of the third and fourth fully cut rings. Emphasis should be placed on reinforcing the soil near the top and waist of the second to fifth rings. The findings demonstrate that the application of flat-face cutterhead in mechanical construction of cross passages is safe, reliable, and efficient, and can provide valuable suggestions for further cutting parameters and soil reinforcement as well. Full article
(This article belongs to the Section Civil Engineering)
Show Figures

Figure 1

Figure 1
<p>Machine of ring cutting and push-in machine.</p>
Full article ">Figure 2
<p>Overview of the cross passage and main tunnel.</p>
Full article ">Figure 3
<p>Layout and dimensions of tunnel segment (dimensions in mm).</p>
Full article ">Figure 4
<p>Finite element analysis model.</p>
Full article ">Figure 5
<p>Schematic diagram of cutting side inner steel plate.</p>
Full article ">Figure 6
<p>Flat-face cutterhead for opening of the cross passage portal.</p>
Full article ">Figure 7
<p>Schematic diagram of the starting stage.</p>
Full article ">Figure 8
<p>Schematic of external loading on main tunnel segments.</p>
Full article ">Figure 9
<p>Schematic diagram of tunnel portal in different stages.</p>
Full article ">Figure 10
<p>Schematic diagram of fully cut and semi-cut segment ring.</p>
Full article ">Figure 11
<p>Horizontal displacement diagram of fully cut ring segment with flat-face cutter head.</p>
Full article ">Figure 12
<p>Vertical displacement diagram of fully cut ring segment with flat-face cutter head.</p>
Full article ">Figure 13
<p>Horizontal displacement diagram of semi-cut ring segment with the flat cutter head.</p>
Full article ">Figure 14
<p>Vertical displacement diagram of semi-cutg ring segment with the flat cutter head.</p>
Full article ">Figure 15
<p>The displacement of semi-cut rings and fully cut ring in the direction of 180°.</p>
Full article ">Figure 16
<p>The displacement of fully cut ring and semi-cut ring in the 90° direction.</p>
Full article ">Figure 17
<p>The displacement of fully cut ring and semi-cut ring in the direction of 270°.</p>
Full article ">Figure 18
<p>Schematic diagram of measuring points.</p>
Full article ">Figure 19
<p>Displacement variation in measuring point in X direction.</p>
Full article ">Figure 20
<p>Displacement variation in measuring points in the Y direction.</p>
Full article ">Figure 21
<p>Displacement variation at the top of the tunnel in the Y direction.</p>
Full article ">Figure 22
<p>Displacement variation at the bottom of the tunnel in the Y direction.</p>
Full article ">Figure 23
<p>The displacement variation in the front of the main tunnel in the x direction.</p>
Full article ">Figure 24
<p>The displacement variation in the back of the main tunnel in the x direction.</p>
Full article ">Figure 25
<p>Side view of segment ring angle selection.</p>
Full article ">Figure 26
<p>Axial force distribution of fully cut ring at different construction stages.</p>
Full article ">Figure 27
<p>Axial force distribution of semi-cut ring tube at different construction stages.</p>
Full article ">Figure 28
<p>Bending moment diagram of the fully cut ring at different construction stages (KN·M).</p>
Full article ">Figure 29
<p>Bending moment diagram of semi-cut ring at different construction stages (KN·M).</p>
Full article ">
Back to TopTop