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

Search Results (2,309)

Search Parameters:
Keywords = seismic performance

Order results
Result details
Results per page
Select all
Export citation of selected articles as:
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 86
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
Show Figures

Figure 1

Figure 1
<p>Details of tested specimens: (<b>a</b>) F60S3U; (<b>b</b>) F60S3SB and F60S3SBD.</p>
Full article ">Figure 2
<p>Used reinforcements in the study: (<b>a</b>) SBPDN1275/1420 bar (U); (<b>b</b>) SBPDN1080/1230 bar (USD).</p>
Full article ">Figure 3
<p>Test methods: (<b>a</b>) test setup; (<b>b</b>) LVDT position.</p>
Full article ">Figure 4
<p>Loading protocols.</p>
Full article ">Figure 5
<p>Damage development and failure of the tested columns (Blue: pull direction, Red: push direction).</p>
Full article ">Figure 6
<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>
Full article ">Figure 7
<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>
Full article ">Figure 8
<p>Comparison of the envelope curves of the specimens: (<b>a</b>) effect of loading method; (<b>b</b>) effect of unbonded length.</p>
Full article ">Figure 9
<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>
Full article ">Figure 10
<p>Calculation method of equivalent viscous damping coefficient.</p>
Full article ">Figure 11
<p>Equivalent viscous damping coefficients of tested specimens.</p>
Full article ">Figure 12
<p>Column modeling and analysis method.</p>
Full article ">Figure 13
<p>Stress–strain model of concrete.</p>
Full article ">Figure 14
<p>Stress–strain model of steel rebar.</p>
Full article ">Figure 15
<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>
Full article ">Figure 16
<p>(<b>a</b>–<b>c</b>) Comparison of skeleton curve between experimental and simulated results.</p>
Full article ">Figure 17
<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>
Full article ">Figure 18
<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>
Full article ">Figure 19
<p>Comparison between experimental and simulated curves: (<b>a</b>) SB columns, (<b>b</b>) SBD columns.</p>
Full article ">Figure 20
<p>Strength degradation ratio vs. the number of loading cycles.</p>
Full article ">Figure 21
<p>Relationship between frequency and degradation at drift ratio 2%.</p>
Full article ">Figure 22
<p>Comparison between experimental and adjusted curves: (<b>a</b>) SB columns, (<b>b</b>) SBD columns.</p>
Full article ">Figure 23
<p>Proposal of a simplified equivalent viscous damping coefficient model.</p>
Full article ">Figure 24
<p>Comparison between the test and analysis results: (<b>a</b>) MRC specimens, (<b>b</b>) NC specimens.</p>
Full article ">
18 pages, 10228 KiB  
Article
Seismic Fragility Analysis of the Transmission Tower-Line System Considering Bolt Slippage
by Jia-Xiang Li, Chao Zhang and Jin-Peng Cheng
Appl. Sci. 2024, 14(23), 10909; https://doi.org/10.3390/app142310909 - 25 Nov 2024
Viewed by 120
Abstract
Seismic fragility analysis can directly reflect the damage probability of the transmission tower-line system and effectively evaluate the seismic performance. To accurately estimate the failure probability of transmission tower-line systems under earthquakes, the tower-line system model considering bolt slippage was established in this [...] Read more.
Seismic fragility analysis can directly reflect the damage probability of the transmission tower-line system and effectively evaluate the seismic performance. To accurately estimate the failure probability of transmission tower-line systems under earthquakes, the tower-line system model considering bolt slippage was established in this paper. Subsequently, through the pushover analysis, different limit states of transmission lines subjected to earthquakes were determined. Finally, 15 seismic waves were selected to study the seismic fragility of a tower-line system by conducting an incremental dynamic analysis on the tower-line system, and seismic fragility curves were drawn. The influences of bolt slippage and foundation deformation on the seismic fragility of the tower are discussed. The results show that the bolt slippage behavior can affect the seismic fragility of tower-line systems and has opposite effects on the transmission tower with a fixed foundation and the transmission tower with foundation deformation. Full article
(This article belongs to the Section Civil Engineering)
Show Figures

Figure 1

Figure 1
<p>Material model.</p>
Full article ">Figure 2
<p>The skeleton curve of the bolt joint.</p>
Full article ">Figure 3
<p>The FEM of the tower-line system.</p>
Full article ">Figure 4
<p>Acceleration response spectrum comparison.</p>
Full article ">Figure 5
<p>Transmission tower section division.</p>
Full article ">Figure 6
<p>Base shear–tower top displacement curves.</p>
Full article ">Figure 7
<p>The maximum stress ratio of the transmission tower: (<b>a</b>) main members of RM; (<b>b</b>) main members of JM; (<b>c</b>) diagonal members of RM; (<b>d</b>) diagonal members of JM.</p>
Full article ">Figure 7 Cont.
<p>The maximum stress ratio of the transmission tower: (<b>a</b>) main members of RM; (<b>b</b>) main members of JM; (<b>c</b>) diagonal members of RM; (<b>d</b>) diagonal members of JM.</p>
Full article ">Figure 8
<p>The failure members of the tower: (<b>a</b>) RM; (<b>b</b>) JM.</p>
Full article ">Figure 9
<p>IDA curves for TTLSs: (<b>a</b>) RM; (<b>b</b>) JM.</p>
Full article ">Figure 10
<p>Calculation models of seismic demand index: (<b>a</b>) RM; (<b>b</b>) JM.</p>
Full article ">Figure 11
<p>Seismic fragility curves of the TTLS with fixed foundation.</p>
Full article ">Figure 12
<p>Layered schematic diagram of the transmission tower.</p>
Full article ">Figure 13
<p><span class="html-italic">ISDR</span> of transmission tower: (<b>a</b>) RM; (<b>b</b>) JM.</p>
Full article ">Figure 14
<p>Node number at the leg of tower 1.</p>
Full article ">Figure 15
<p>IDA curves of TTLS: (<b>a</b>) RM; (<b>b</b>) JM.</p>
Full article ">Figure 16
<p>Calculation models of seismic demand index: (<b>a</b>) RM; (<b>b</b>) JM.</p>
Full article ">Figure 17
<p>Seismic fragility curves of the TTLS with foundation settlement.</p>
Full article ">Figure 18
<p><span class="html-italic">ISDR</span> of transmission tower: (<b>a</b>) RM; (<b>b</b>) JM.</p>
Full article ">Figure 19
<p>IDA curves of TTLSs: (<b>a</b>) RM; (<b>b</b>) JM.</p>
Full article ">Figure 20
<p>Data fitting curves: (a) RM; (b) JM.</p>
Full article ">Figure 21
<p>Seismic fragility curves.</p>
Full article ">Figure 22
<p><span class="html-italic">ISDR</span> values for TTLSs with foundation horizontal displacement: (<b>a</b>) RM; (<b>b</b>) JM.</p>
Full article ">
20 pages, 6919 KiB  
Article
Analysis of Dynamic Characteristics and Seismic Response of Chen Xiang Pavilion in Xi’an Considering the Lower Stylobate
by Kang Liu, Huifang Liao, Bowen Xue, Chenwei Wu, Jianyang Xue, Dejun Song and Hao Xue
Buildings 2024, 14(12), 3742; https://doi.org/10.3390/buildings14123742 - 24 Nov 2024
Viewed by 307
Abstract
This paper presents the dynamic characteristics and seismic performance of the Chen Xiang Pavilion in Xi’an and the influence of the lower stylobate on the dynamic response of the upper wooden structure. An in situ dynamic test was conducted under ambient vibration to [...] Read more.
This paper presents the dynamic characteristics and seismic performance of the Chen Xiang Pavilion in Xi’an and the influence of the lower stylobate on the dynamic response of the upper wooden structure. An in situ dynamic test was conducted under ambient vibration to detect the natural frequencies and vibration modes of the structure. Three numerical models, including the upper wooden structure, the lower stylobate, and the whole structure (wooden structure and stylobate), were established. Dynamic characteristic and seismic response analyses were performed on the calculated models to investigate the influence of the lower stylobate on the dynamic response of the upper wooden structure. The simulation results indicated that the lower stylobate significantly affected the dynamic characteristics of the upper wooden structure above the third order. The seismic responses of the upper wooden structure were amplified because of the lower stylobate. Under different excitations, the displacement response of the whole structure was up to 1.99 times relative to the upper wooden structure, and the structural shear forces were increased by 15.3%. The dynamic amplification coefficient was magnified from 0.742~0.948 to 1.024~1.776. The Chen Xiang Pavilion has a good energy dissipation capacity, but the lower stylobate is unfavorable for its earthquake resistance. Full article
Show Figures

Figure 1

Figure 1
<p>Photo of Chen Xiang Pavilion in Xi’an.</p>
Full article ">Figure 2
<p>Elevation, sectional, and plan views of Chen Xiang Pavilion.</p>
Full article ">Figure 3
<p>Flow chart of data acquisition.</p>
Full article ">Figure 3 Cont.
<p>Flow chart of data acquisition.</p>
Full article ">Figure 4
<p>Time history response curve of eave column top in EW direction.</p>
Full article ">Figure 5
<p>Power spectral density–frequency curves of the tops of the columns in the NS and EW directions.</p>
Full article ">Figure 6
<p>First two translation modes of the upper wooden structure in the NS and EW directions.</p>
Full article ">Figure 7
<p>Finite element models of the upper wooden structure, the lower stylobate, and the whole structure.</p>
Full article ">Figure 8
<p>First vibration modes of the calculation model in the NS and EW directions.</p>
Full article ">Figure 9
<p>First ten-order natural frequencies of three models.</p>
Full article ">Figure 10
<p>Several vibration modes of the upper wooden structure and the whole structure.</p>
Full article ">Figure 11
<p>Comparison between seismic response spectrum and target response spectrum.</p>
Full article ">Figure 12
<p>Peak acceleration response.</p>
Full article ">Figure 13
<p>Maximum inter-structural drift angle.</p>
Full article ">Figure 14
<p>Maximum shear force of layer under El Centro wave.</p>
Full article ">
21 pages, 1884 KiB  
Article
Seismic Analysis of Non-Regular Structures Based on Fragility Curves: A Study on Reinforced Concrete Frames
by Giovanni Smiroldo, Marco Fasan and Chiara Bedon
Buildings 2024, 14(12), 3734; https://doi.org/10.3390/buildings14123734 - 23 Nov 2024
Viewed by 395
Abstract
The seismic performance and expected structural damage in reinforced concrete (RC) frames, as in many others, is a critical aspect for design. In this study, a set of RC frames characterized by increasing in-plan and in-height non-regularity is specifically investigated. Four code-conforming three-dimensional [...] Read more.
The seismic performance and expected structural damage in reinforced concrete (RC) frames, as in many others, is a critical aspect for design. In this study, a set of RC frames characterized by increasing in-plan and in-height non-regularity is specifically investigated. Four code-conforming three-dimensional (3D) buildings with varying regularity levels are numerically analyzed. Their seismic assessment is conducted by using unscaled real ground motion records (61 in total) and employing non-linear dynamic simulations within the Cloud Analysis framework. Three distinct intensity measures (IMs) are used to evaluate the impact of structural non-regularity on their seismic performance. Furthermore, fragility curves are preliminary derived based on conventional linear regression models and lognormal distribution. In contrast with the initial expectations and the typical results of non-linear dynamic analyses, the presented comparative results of the fragility curves show that the non-regularity level increase for the examined RC frames does not lead to progressively increasing fragility. Upon these considerations on the initial numerical findings, a re-evaluation of the methodology is performed using a reduced subset of ground motion records, in order to account for potential biases in their selection. Moreover, to uncover deeper insights into the unexpected outcomes, a logistic regression based on a maximum likelihood estimate is also employed to develop fragility curves. Comparative results are thus critically discussed, showing that the herein considered fragility development methods may lead to seismic assessment outcomes for code-conforming non-regular buildings that are in contrast with those of raw structural analyses. In fact, the considered building code design provisions seem to compensate non-regularity-induced torsional effects. Full article
(This article belongs to the Collection Advanced Concrete Structures in Civil Engineering)
26 pages, 834 KiB  
Article
Probabilistic Evaluation Method of Wind Resistance of Membrane Roofs Based on Aerodynamic Stability
by Weiju Song, Hongbo Liu and Heding Yu
Buildings 2024, 14(12), 3725; https://doi.org/10.3390/buildings14123725 - 22 Nov 2024
Viewed by 162
Abstract
The membrane structure or membrane roofing system is lightweight and flexible, with wind being the primary cause of structural and membrane material failure. To evaluate the disaster prevention and mitigation capacity of the membrane roofing system and enhance the wind disaster risk management [...] Read more.
The membrane structure or membrane roofing system is lightweight and flexible, with wind being the primary cause of structural and membrane material failure. To evaluate the disaster prevention and mitigation capacity of the membrane roofing system and enhance the wind disaster risk management capabilities, this paper studies the exceedance probability evaluation method for different wind resistance requirements of membrane roofs. Taking Hangzhou in China as an example, the design wind speed risk curve fitted by polynomial is obtained by referring to the PEER performance-based seismic design method and considering the randomness of the wind field. A polynomial fitting method is employed to obtain the design wind speed hazard curve. Considering the nonlinear characteristics of the membrane roof structure, the relationship between the roof’s wind resistance requirements (vertical displacement limits) and wind speed spectrum values is approximated using a power function. An annual average exceedance probability expression is derived for different normal deformation demand values of the membrane roofs under wind load. Based on this, a wind resistance probability evaluation method for membrane roofs considering aerodynamic stability is proposed, along with specific steps and related analytical formulas. The results indicate that polynomial fitting provides an effective simplification for deriving the annual average exceedance probability expression for the wind resistance demand of membrane roofs. The performance-based wind resistance probability evaluation method allows for obtaining exceedance probability values for different displacement requirements with minimal structural analysis, which enriches the wind resistance design theory of membrane roofs and further ensures the structural safety of tension membrane roofs under wind load. Full article
18 pages, 7320 KiB  
Article
Direct Tensile Test Method for Shotcrete
by Oleg V. Kabancev and Oleg A. Simakov
Buildings 2024, 14(12), 3713; https://doi.org/10.3390/buildings14123713 - 21 Nov 2024
Viewed by 221
Abstract
This study substantiates the need for direct tensile strength testing of shotcrete and fiber-reinforced shotcrete, rather than relying on indirect methods, to accurately reflect material performance under biaxial stress conditions when used for structural reinforcement. Experiments on field specimens confirmed that tensile strength [...] Read more.
This study substantiates the need for direct tensile strength testing of shotcrete and fiber-reinforced shotcrete, rather than relying on indirect methods, to accurately reflect material performance under biaxial stress conditions when used for structural reinforcement. Experiments on field specimens confirmed that tensile strength values derived through direct testing differ significantly from those calculated based on compressive strength. The study presents a new testing methodology with optimized specimen dimensions (32, 40, 50, and 82 mm diameter cylinders with length-to-diameter ratios of 3.0) to mitigate eccentricity effects, ensuring normal-section failure. Results show that tensile strength values for fiber-reinforced shotcrete with brass-coated fibers (13–15 mm length, 0.3–0.5 mm diameter, 30 kg/m3 dosage) reached 68 MPa, compared to 60 MPa for standard shotcrete, while basalt-fiber reinforcement (6 mm length, 1% by weight) resulted in 42 MPa. The initial modulus of elasticity for unreinforced shotcrete was 280 × 103 MPa, with fiber reinforcement slightly increasing this value to 287 × 103 MPa. The findings support a direct approach to testing, providing a foundation for developing predictive methodologies for fiber-reinforced shotcrete properties based on reinforcement type and dosage. These results are essential for applications such as seismic strengthening, where accurate tensile characteristics are critical for performance under dynamic loading. Full article
(This article belongs to the Special Issue Safety and Optimization of Building Structures—2nd Edition)
Show Figures

Figure 1

Figure 1
<p>Causes requiring restoration of reinforced concrete and masonry structures.</p>
Full article ">Figure 2
<p>Damage to exterior masonry walls due to seismic impacts.</p>
Full article ">Figure 3
<p>Sample of masonry structures reinforced with a one-sided shotcrete overlay.</p>
Full article ">Figure 4
<p>Results of the study on failure mechanisms at the interface between shotcrete overlays and masonry.</p>
Full article ">Figure 5
<p>Model of interaction between structural elements with compromised internal bonds and the external reinforcing element: under compression and shear conditions (<b>a</b>), tension and shear conditions (<b>b</b>).</p>
Full article ">Figure 6
<p>Test scheme for notched beam specimens, according to EN 14651 [<a href="#B16-buildings-14-03713" class="html-bibr">16</a>].</p>
Full article ">Figure 7
<p>Test scheme for notched beam specimens, according to ASTM C1609 [<a href="#B17-buildings-14-03713" class="html-bibr">17</a>].</p>
Full article ">Figure 8
<p>Test scheme for panel specimens, according to ASTM C1550 [<a href="#B18-buildings-14-03713" class="html-bibr">18</a>].</p>
Full article ">Figure 9
<p>Shotcrete and fiber-reinforced shotcrete specimens for testing with diameters of 64 and 72 mm; (<b>a</b>) extracted from panels, (<b>b</b>) cast in plastic cylinder molds with diameters of 32 mm, 40 mm, 50 mm, and 82 mm.</p>
Full article ">Figure 10
<p>General view of the test series of specimens for compressive strength testing.</p>
Full article ">Figure 11
<p>General view of the test series of specimens for tensile strength testing.</p>
Full article ">Figure 12
<p>Testing of control specimens: (<b>a</b>) compression; (<b>b</b>) tension.</p>
Full article ">Figure 13
<p>Support “cups” for gripping specimens in tensile testing.</p>
Full article ">Figure 14
<p>Tests for determining the modulus of elasticity in compression.</p>
Full article ">Figure 15
<p>Tensile tests on specimens with diameters of 32 and 40 mm.</p>
Full article ">Figure 16
<p>Tensile tests on specimens with diameters of 50 and 80 mm. (<b>a</b>) formation of rectilinear inclined crack, (<b>b</b>) formation of nonlinear crack with displacement.</p>
Full article ">Figure 17
<p>Specimen failure pattern.</p>
Full article ">
22 pages, 10421 KiB  
Article
Distributed High-Speed Videogrammetry for Real-Time 3D Displacement Monitoring of Large Structure on Shaking Table
by Haibo Shi, Peng Chen, Xianglei Liu, Zhonghua Hong, Zhen Ye, Yi Gao, Ziqi Liu and Xiaohua Tong
Remote Sens. 2024, 16(23), 4345; https://doi.org/10.3390/rs16234345 - 21 Nov 2024
Viewed by 248
Abstract
The accurate and timely acquisition of high-frequency three-dimensional (3D) displacement responses of large structures is crucial for evaluating their condition during seismic excitation on shaking tables. This paper presents a distributed high-speed videogrammetric method designed to rapidly measure the 3D displacement of large [...] Read more.
The accurate and timely acquisition of high-frequency three-dimensional (3D) displacement responses of large structures is crucial for evaluating their condition during seismic excitation on shaking tables. This paper presents a distributed high-speed videogrammetric method designed to rapidly measure the 3D displacement of large shaking table structures at high sampling frequencies. The method uses non-coded circular targets affixed to key points on the structure and an automatic correspondence approach to efficiently estimate the extrinsic parameters of multiple cameras with large fields of view. This process eliminates the need for large calibration boards or manual visual adjustments. A distributed computation and reconstruction strategy, employing the alternating direction method of multipliers, enables the global reconstruction of time-sequenced 3D coordinates for all points of interest across multiple devices simultaneously. The accuracy and efficiency of this method were validated through comparisons with total stations, contact sensors, and conventional approaches in shaking table tests involving large structures with RCBs. Additionally, the proposed method demonstrated a speed increase of at least six times compared to the advanced commercial photogrammetric software. It could acquire 3D displacement responses of large structures at high sampling frequencies in real time without requiring a high-performance computing cluster. Full article
Show Figures

Graphical abstract

Graphical abstract
Full article ">Figure 1
<p>Framework of the proposed videogrammetric method.</p>
Full article ">Figure 2
<p>General distributed videogrammetric network.</p>
Full article ">Figure 3
<p>Stereo-matching method of circular targets in large FOV (red dots indicate SIFT feature points of stereo images).</p>
Full article ">Figure 4
<p>Distributed computation and reconstruction strategy.</p>
Full article ">Figure 5
<p>(<b>a</b>) Real structure model. (<b>b</b>) Camera layout and spatial coordinate system. (<b>c</b>) Measurement point distribution.</p>
Full article ">Figure 6
<p>Measurement errors between the videogrammetry and the total station at each checkpoint in the X, Y, and Z directions.</p>
Full article ">Figure 7
<p>Three-dimensional positioning errors of the checkpoint calculated using different methods after each seismic wave load.</p>
Full article ">Figure 8
<p>Comparison of displacement and acceleration response histories obtained by the proposed videogrammetry and contact sensors at points <span class="html-italic">R</span><sub>3</sub> and <span class="html-italic">R</span><sub>18</sub> subjected to different seismic excitations: (<b>a</b>) Experiment No. 1; (<b>b</b>) Experiment No. 3; (<b>c</b>) Experiment No. 5.</p>
Full article ">Figure 9
<p>Time consumption and mean reprojection error of different methods for reconstructing the shaking table dataset.</p>
Full article ">Figure 10
<p>Time consumption of different methods for reconstructing the shaking table dataset.</p>
Full article ">Figure 11
<p>Three-dimensional displacement response histories of measurement points distributed across the coupling beams during (<b>a</b>) Experiment No. 1, (<b>b</b>) Experiment No. 3, and (<b>c</b>) Experiment No. 5.</p>
Full article ">
19 pages, 16335 KiB  
Article
The Controlling Factors and Prediction of Deep-Water Mass Transport Deposits in the Pliocene Qiongdongnan Basin, South China Sea
by Jiawang Ge, Xiaoming Zhao, Qi Fan, Weixin Pang, Chong Yue and Yueyao Chen
J. Mar. Sci. Eng. 2024, 12(12), 2115; https://doi.org/10.3390/jmse12122115 - 21 Nov 2024
Viewed by 268
Abstract
Large-scaled submarine slides or mass transport deposits (MTDs) widely occurred in the Pliocene Qiongdongnan Basin, South China Sea. The good seismic mapping and distinctive topography, as well as the along-striking variation in sediment supply, make it an ideal object to explore the linkage [...] Read more.
Large-scaled submarine slides or mass transport deposits (MTDs) widely occurred in the Pliocene Qiongdongnan Basin, South China Sea. The good seismic mapping and distinctive topography, as well as the along-striking variation in sediment supply, make it an ideal object to explore the linkage of controlling factors and MTD distribution. The evaluation of the main controlling factors of mass transport deposits utilizes the analysis of terrestrial catastrophes as a reference based on the GIS-10.2 software. The steepened topography is assumed to be an external influence on triggering MTDs; therefore, the MTDs are mapped to the bottom interface of the corresponding topography strata. Based on detailed seismic and well-based observations from multiple phases of MTDs in the Pliocene Qiongdongnan Basin (QDNB), the interpreted controlling factors are summarized. Topographic, sedimentary, and climatic factors are assigned to the smallest grid cell of this study. Detailed procedures, including correlation analysis, significance check, and recursive feature elimination, are conducted. A random forest artificial intelligence algorithm was established. The mean value of the squared residuals of the model was 0.043, and the fitting degree was 82.52. To test the stability and accuracy of this model, the training model was used to calibrate the test set, and five times 2-fold cross-validation was performed. The area under the curve mean value is 0.9849, indicating that the model was effective and stable. The most related factors are correlated to the elevation, flow direction, and slope gradient. The predicted results were consistent with the seismic interpretation results. Our study indicates that a random forest artificial intelligence algorithm could be useful in predicting the susceptibility of deep-water MTDs and can be applied to other study areas to predict and avoid submarine disasters caused by wasting processes. Full article
Show Figures

Figure 1

Figure 1
<p>Location of QDNB in the South China Sea (<b>a</b>) and tectonic unit as well as the well location of the study area (<b>b</b>). The black line indicates six secondary blocks named from 1 to 6 in the study area from west to east.</p>
Full article ">Figure 2
<p>Elevation maps of the main surfaces, including T30 (<b>a</b>), T29 (<b>b</b>), T28 (<b>c</b>), T27 (<b>d</b>), and T20 (<b>e</b>) of the Pliocene Yinggehai Formation in the Qiongdongnan Basin, South China Sea.</p>
Full article ">Figure 3
<p>Seismic profiles 3, 10, and AA’ showing the MTDs identified from T30 to T29. TWT: two-way travel time; MTDs: mass-transport deposits. See locations in <a href="#jmse-12-02115-f001" class="html-fig">Figure 1</a>b.</p>
Full article ">Figure 3 Cont.
<p>Seismic profiles 3, 10, and AA’ showing the MTDs identified from T30 to T29. TWT: two-way travel time; MTDs: mass-transport deposits. See locations in <a href="#jmse-12-02115-f001" class="html-fig">Figure 1</a>b.</p>
Full article ">Figure 4
<p>MTD plane distribution map in four sequences ((<b>a</b>) T30-T29, (<b>b</b>) T29-T28, (<b>c</b>) T28-T27, (<b>d</b>) T27-T20) of Yinggehai Formation in Qiongdongnan Basin, South China Sea.</p>
Full article ">Figure 5
<p>Schematic diagram of morphological characteristics and geometric parameters of slope topography [<a href="#B53-jmse-12-02115" class="html-bibr">53</a>].</p>
Full article ">Figure 6
<p>Digital flow chart of MTD controlling factors. The numerical process of MTD controlling factors is described in detail.</p>
Full article ">Figure 7
<p>The specific steps of random forest. Describe the algorithm flow of random forest in detail.</p>
Full article ">Figure 8
<p>Correlation and visibility of MTD trigger factors. a: Elevation; b: Slope gradient; c: Profle curvature; d: Accretion rate; e: Slope direction; f: flow direction; g: Sediment fluxes; h: Progradation rate; i: Ratio of accretion to progradation; j: Sea level fluctuation rate; k: flow rate; l: Plane curvature.</p>
Full article ">Figure 9
<p>The recursive feature elimination root mean square error trend diagram of MTD controlling factors.</p>
Full article ">Figure 10
<p>The importance ranking of MTD controlling factors. The controlling factors are ranked by importance analysis. Notice that the elevation, flow direction, and slope gradient rank high.</p>
Full article ">Figure 11
<p>Random forest model receiver operating characteristics reflected the high accuracy of the model.</p>
Full article ">Figure 12
<p>MTD prediction results in four third-order sequences ((<b>a</b>) T30-T29; (<b>b</b>) T29-T28; (<b>c</b>) T28-T27; (<b>d</b>) T27-T20) of Pliocene in QDNB, South China Sea. Notice that the results were consistent with the facts presented in <a href="#jmse-12-02115-f004" class="html-fig">Figure 4</a>.</p>
Full article ">Figure 13
<p>Partial dependency map of MTD controlling factors: Eleveation (<b>a</b>), Flow direction (<b>b</b>), Slope gradient (<b>c</b>), Ratio of accretion to progration (<b>d</b>), Progradation rate (<b>e</b>), Flow rate (<b>f</b>), Profile curvature (<b>g</b>), Plane curvature (<b>h</b>), Accretion rate (<b>i</b>) and Sea level fluctuation rate (<b>j</b>).</p>
Full article ">
28 pages, 9654 KiB  
Article
Time Series Foundation Models and Deep Learning Architectures for Earthquake Temporal and Spatial Nowcasting
by Alireza Jafari, Geoffrey Fox, John B. Rundle, Andrea Donnellan and Lisa Grant Ludwig
GeoHazards 2024, 5(4), 1247-1274; https://doi.org/10.3390/geohazards5040059 - 21 Nov 2024
Viewed by 273
Abstract
Advancing the capabilities of earthquake nowcasting, the real-time forecasting of seismic activities, remains crucial for reducing casualties. This multifaceted challenge has recently gained attention within the deep learning domain, facilitated by the availability of extensive earthquake datasets. Despite significant advancements, the existing literature [...] Read more.
Advancing the capabilities of earthquake nowcasting, the real-time forecasting of seismic activities, remains crucial for reducing casualties. This multifaceted challenge has recently gained attention within the deep learning domain, facilitated by the availability of extensive earthquake datasets. Despite significant advancements, the existing literature on earthquake nowcasting lacks comprehensive evaluations of pre-trained foundation models and modern deep learning architectures; each focuses on a different aspect of data, such as spatial relationships, temporal patterns, and multi-scale dependencies. This paper addresses the mentioned gap by analyzing different architectures and introducing two innovative approaches called Multi Foundation Quake and GNNCoder. We formulate earthquake nowcasting as a time series forecasting problem for the next 14 days within 0.1-degree spatial bins in Southern California. Earthquake time series are generated using the logarithm energy released by quakes, spanning 1986 to 2024. Our comprehensive evaluations demonstrate that our introduced models outperform other custom architectures by effectively capturing temporal-spatial relationships inherent in seismic data. The performance of existing foundation models varies significantly based on the pre-training datasets, emphasizing the need for careful dataset selection. However, we introduce a novel method, Multi Foundation Quake, that achieves the best overall performance by combining a bespoke pattern with Foundation model results handled as auxiliary streams. Full article
Show Figures

Figure 1

Figure 1
<p>Illustration of the construction of a nowcast model for California. The nowcast is a 2-parameter filter on the small earthquake seismicity [<a href="#B42-geohazards-05-00059" class="html-bibr">42</a>,<a href="#B43-geohazards-05-00059" class="html-bibr">43</a>]. (<b>a</b>) Seismicity in the Los Angeles region since 1960, M &gt; 3.29. (<b>b</b>) Monthly rate of small earthquakes as cyan vertical bars. The blue curve is the 36-month exponential moving average (EMA). (<b>c</b>) Mean rate of small earthquakes since 1970. (<b>d</b>) Nowcast curve that is the result of applying the optimized EMA and corrections for the time-varying small earthquake rate to the small earthquake seismicity. (<b>e</b>) Optimized receiver operating characteristic (ROC) curve (red line) used in the machine learning algorithm. Skill is the area under the ROC curve and is used in the optimization. Skill trade-off diagram shows the range of models used in the optimization.</p>
Full article ">Figure 2
<p>Image showing the application of the trained QuakeGPT transformer to an independent, scaled nowcast validation curve (green shading), followed by prediction of future values beyond the end of the nowcast curve (magenta shading). In this model, 36 previous values are used to predict the next value. Dots show the predictions and the solid line shows the nowcast curve whose values are to be predicted. Green dots show the predictions of the transformer up to the last 37 values. The 36 blue dots are predictions that were made and then fed back into the transformer to predict the final point (red dot). In this model, 50 members of an ensemble of runs were used to make the predictions. The dots represent the mean predictions. Brown areas represent the 1-sigma standard deviations to the mean values. In this model, 2021 years of simulation data were used to train the model.</p>
Full article ">Figure 3
<p>Distribution of earthquake epicenters in Southern California (32° N to 36° N, −120° to −114°) from USGS data (1986–2024). The scatter plot shows the spatial density of seismic events used to analyze and optimize spatial bins for earthquake nowcasting.</p>
Full article ">Figure 4
<p>The 500 most active and vulnerable spatial bins, marked in blue, selected for analysis out of the total 2400, based on the frequency of earthquakes from 1986 to 2024. This selection focuses on high-risk areas.</p>
Full article ">Figure 5
<p>Six time series from randomly selected spatial bins, highlighting earthquakes of magnitude greater than 5.</p>
Full article ">Figure 6
<p>The final graph structure representing the 500 most active bins, created using an epsilon of 0.15 degrees. Initially forming a multi-component graph, components are linked to ensure full connectivity.</p>
Full article ">Figure 7
<p>Released energy time series plots for six randomly selected spatial bins, comparing model predictions (GNNCoder one-layer, DilatedRNN, TiDE, iTransformer-M4) against actual observed seismic activities. The brown line represents our GNN approach, which shows a closer match with the actual time series, capturing crucial upward slopes that may signal an impending earthquake. The green and red lines occasionally miss these trends, making more errors where even slight changes in seismic activity are critical. The purple line from the iTransformer-M4 model fails to accurately capture the time series values and exhibits excessive fluctuations.</p>
Full article ">Figure 8
<p>This plot illustrates the spatial bins overlaid on the fault lines to assess the extent to which the fault lines are captured by the bins (graph nodes). It highlights the limitations of the current graph, where some critical fault lines fall outside the spatial bins, impacting the performance of deeper GNN models like the GNNCoder 3-layer model.</p>
Full article ">
14 pages, 4536 KiB  
Article
Numerical Simulation of Seismoacoustic Wave Transformation at Sea–Land Interface
by Grigory Dolgikh, Mikhail Bolsunovskii, Denis Zharkov, Ruslan Zhostkov, Dmitriy Presnov, Andrey Razin and Andrey Shurup
J. Mar. Sci. Eng. 2024, 12(12), 2112; https://doi.org/10.3390/jmse12122112 - 21 Nov 2024
Viewed by 278
Abstract
This study considers seismoacoustic wave propagation through the land–sea interface, i.e., in the presence of a coastal wedge, taking into account the real bottom bathymetry. It is of interest in the problems of coastal monitoring and environmental studies. An effective numerical model based [...] Read more.
This study considers seismoacoustic wave propagation through the land–sea interface, i.e., in the presence of a coastal wedge, taking into account the real bottom bathymetry. It is of interest in the problems of coastal monitoring and environmental studies. An effective numerical model based on the finite element method is proposed and implemented. An approximate analytical solution in the fluid and an asymptotic analytical solution for the surface seismic wave on the shore are considered to validate the numerical model. It is shown that in field experiment conditions the hydroacoustic signal generated by an underwater source with a power of ~200 W is transformed into a seismic wave on the shore with an amplitude of units of nanometers at distances of several kilometers, which can be measured by a sensitive sensor. An extensive series of numerical simulations with different model parameters was performed, which allowed us to evaluate the most appropriate propagation medium parameters to match the observed and calculated data. Full article
Show Figures

Figure 1

Figure 1
<p>Map of the measurement site (Posieta Bay, Gamova Peninsula, near the Marine Experimental Station of V.I. Il’ichev Pacific Oceanological Institute). S1, S2—source points; LS—coastal laser strainmeter (receiver point).</p>
Full article ">Figure 2
<p>Realistic model geometry and finite element mesh. I—source; II—area of averaging displacement amplitude values (receiver); III—perfectly matched layers (PMLs).</p>
Full article ">Figure 3
<p>Geometry of the ASA coastal wedge model used for verification and validation.</p>
Full article ">Figure 4
<p>Pressure values at point (<span class="html-italic">x</span> = 2000 m, <span class="html-italic">y</span> = −30 m) as a function of the quality factor (KQ). The results show the stabilization of the solution as the KQ reaches 1, with deviations under 0.1%.</p>
Full article ">Figure 5
<p>Dependence of the pressure value at the point <span class="html-italic">x</span> = 2000 m, <span class="html-italic">y</span> = −30 m of the Δ—relative changes in the input parameters.</p>
Full article ">Figure 6
<p>Comparison of numerical simulation results and analytical solutions for hydroacoustic signal transmission loss (<b>a</b>) and vertical profile of the amplitude of displacement components on land normalized to the value on the surface (<b>b</b>).</p>
Full article ">Figure 7
<p>Propagation of a pulsed signal over time, time stamps: (<b>a</b>)—2 s; (<b>b</b>)—3 s; (<b>c</b>)—4.7 s; (<b>d</b>)—5.7 s. The color scale is limited to highlight weak variations. The signal, generated at a shallow depth, produces both body waves and a Scholte surface wave at the water–bottom interface.</p>
Full article ">Figure 8
<p>Calculated pulse seismogram at the laser strainmeter location. (<b>a</b>) Vertical displacements and (<b>b</b>) horizontal displacements show the arrival of body waves at 2.5 s intervals, followed by Rayleigh waves at 5 s intervals.</p>
Full article ">Figure 9
<p>Dependence of the residual <math display="inline"><semantics> <mrow> <mrow> <mrow> <mfenced close="|" open="|"> <mrow> <msub> <mi>U</mi> <mi>x</mi> </msub> <mo>−</mo> <msub> <mover accent="true"> <mi>U</mi> <mo stretchy="false">^</mo> </mover> <mi>x</mi> </msub> </mrow> </mfenced> </mrow> <mo>/</mo> <mrow> <msub> <mi>U</mi> <mi>x</mi> </msub> </mrow> </mrow> </mrow> </semantics></math> between the experimentally measured <math display="inline"><semantics> <mrow> <msub> <mi>U</mi> <mi>x</mi> </msub> </mrow> </semantics></math> horizontal displacement and its estimate <math display="inline"><semantics> <mrow> <msub> <mover accent="true"> <mi>U</mi> <mo stretchy="false">^</mo> </mover> <mi>x</mi> </msub> </mrow> </semantics></math>, numerically calculated for radiation Point S1 (<b>a</b>) and Point S2 (<b>b</b>). The dotted oval indicates the region of longitudinal wave velocities <math display="inline"><semantics> <mrow> <msub> <mi>c</mi> <mi>l</mi> </msub> </mrow> </semantics></math>, where good agreement between the experimental and calculated data is observed.</p>
Full article ">Figure 10
<p>Modeling results of displacement amplitudes obtained for the source at Point S1 (<b>a</b>) and at Point S2 (<b>b</b>) (uniform radiation in the band 19–26 Hz with an amplitude of 6.9 ± 1 kPa is assumed). Different colors of the points on the graph correspond to different values of the simulation parameters: 1800 ≤ <math display="inline"><semantics> <mrow> <msub> <mi>c</mi> <mi>l</mi> </msub> </mrow> </semantics></math> ≤ 2000 m/s for Point S1; 2200 ≤ <math display="inline"><semantics> <mrow> <msub> <mi>c</mi> <mi>l</mi> </msub> </mrow> </semantics></math> ≤ 2450 m/s for Point S2 (these ranges are highlighted by dashed ovals in <a href="#jmse-12-02112-f008" class="html-fig">Figure 8</a>).</p>
Full article ">
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 264
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)
Show Figures

Graphical abstract

Graphical abstract
Full article ">Figure 1
<p>Pipeline subjected to longitudinal PGD: (<b>a</b>) 3D view; (<b>b</b>) 2D schematic representation.</p>
Full article ">Figure 2
<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>
Full article ">Figure 3
<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>
Full article ">Figure 4
<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>
Full article ">Figure 5
<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>
Full article ">Figure 6
<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>
Full article ">Figure 7
<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>
Full article ">Figure 8
<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>
Full article ">Figure 9
<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>
Full article ">Figure 10
<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>
Full article ">Figure 11
<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>
Full article ">Figure 12
<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>
Full article ">Figure A1
<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>
Full article ">Figure A2
<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>
Full article ">Figure A3
<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>
Full article ">Figure A4
<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>
Full article ">
23 pages, 10425 KiB  
Article
Hybrid Reinforced Concrete Frames with Engineering Cementitious Composites: Experimental and Numerical Investigations
by Abdulrahman Metawa, Moussa Leblouba and Samer Barakat
Sustainability 2024, 16(22), 10085; https://doi.org/10.3390/su162210085 - 19 Nov 2024
Viewed by 283
Abstract
Reinforced concrete (RC) structures are vulnerable to damage under dynamic loads such as earthquakes, necessitating innovative solutions that enhance both performance and sustainability. This study investigates the integration of Engineered Cementitious Composites (ECC) in RC frames to improve ductility, durability, and energy dissipation [...] Read more.
Reinforced concrete (RC) structures are vulnerable to damage under dynamic loads such as earthquakes, necessitating innovative solutions that enhance both performance and sustainability. This study investigates the integration of Engineered Cementitious Composites (ECC) in RC frames to improve ductility, durability, and energy dissipation while considering cost-effectiveness. To achieve this, the partial replacement of concrete with ECC at key structural locations, such as beam–column joints, was explored through experimental testing and numerical simulations. Small-scale beams with varying ECC replacements were tested for failure modes, load–deflection responses, and crack propagation patterns. Additionally, nonlinear quasi-static cyclic and modal analyses were performed on full RC frames, ECC-reinforced frames, and hybrid frames with ECC at the joints. The results demonstrate that ECC reduces the need for shear reinforcement due to its crack-bridging ability, enhances ductility by up to 25% in cyclic loading scenarios, and lowers the formation of plastic hinges, thereby contributing to improved structural resilience. These findings suggest that ECC is a viable, sustainable solution for achieving resilient infrastructure in seismic regions, with an optimal balance between performance and cost. Full article
(This article belongs to the Special Issue Research Advances in Sustainable Materials and Structural Engineering)
Show Figures

Figure 1

Figure 1
<p>Test specimens: (<b>a</b>) RC-St; (<b>b</b>) RC-ECC; (<b>c</b>) RC-ECC-St; (<b>d</b>) ECC (units: mm).</p>
Full article ">Figure 1 Cont.
<p>Test specimens: (<b>a</b>) RC-St; (<b>b</b>) RC-ECC; (<b>c</b>) RC-ECC-St; (<b>d</b>) ECC (units: mm).</p>
Full article ">Figure 2
<p>Test setup and instrumentation.</p>
Full article ">Figure 3
<p>Beam specimens after testing: (<b>a</b>) RC-St; (<b>b</b>) RC-ECC; (<b>c</b>) RC-ECC-St; (<b>d</b>) ECC.</p>
Full article ">Figure 4
<p>Load–deflection and crack openings: (<b>a</b>) RC; (<b>b</b>) RC-ECC; (<b>c</b>) RC-ECC-St; (<b>d</b>) ECC; (<b>e</b>) idealized load–deflection curves.</p>
Full article ">Figure 4 Cont.
<p>Load–deflection and crack openings: (<b>a</b>) RC; (<b>b</b>) RC-ECC; (<b>c</b>) RC-ECC-St; (<b>d</b>) ECC; (<b>e</b>) idealized load–deflection curves.</p>
Full article ">Figure 5
<p>Stress–strain relationship of confined and unconfined concrete (compression side only).</p>
Full article ">Figure 6
<p>Stress–strain relationship of steel reinforcement.</p>
Full article ">Figure 7
<p>Stress–strain relationship of ECC in compression and tension.</p>
Full article ">Figure 8
<p>RC, EC-ECC, and ECC case study frames.</p>
Full article ">Figure 9
<p>FEMA 461 loading protocol (partial signal).</p>
Full article ">Figure 10
<p>Force–displacement curves of the three frames.</p>
Full article ">Figure 11
<p>Damage locations in the RC frame.</p>
Full article ">Figure 12
<p>Damage locations in the RC–ECC frame.</p>
Full article ">Figure 13
<p>Damage locations in the full ECC frame.</p>
Full article ">Figure 14
<p>Cumulative energy dissipation and residual deflection ratio.</p>
Full article ">
12 pages, 2559 KiB  
Article
Research on Oil and Gas-Bearing Zone Prediction and Identification Based on the SVD–K-Means Algorithm—A Case Study of the WZ6-1 Oil-Bearing Structure in the Beibu Gulf Basin, South China Sea
by Zhilong Chen, Renyi Wang, Biao Xu and Jianghang Zhu
Energies 2024, 17(22), 5771; https://doi.org/10.3390/en17225771 - 19 Nov 2024
Viewed by 344
Abstract
The WZ6-1 oil-bearing structure in the Beibu Gulf Basin of the South China Sea has well-developed faults with significant variations in fault sealing capacity, resulting in a complex and highly variable distribution of oil, gas, and water, and limited understanding of hydrocarbon accumulation [...] Read more.
The WZ6-1 oil-bearing structure in the Beibu Gulf Basin of the South China Sea has well-developed faults with significant variations in fault sealing capacity, resulting in a complex and highly variable distribution of oil, gas, and water, and limited understanding of hydrocarbon accumulation patterns. Traditional methods, such as single seismic attributes and linear fusion of multiple seismic attributes, have proven ineffective in identifying and predicting oil and gas-bearing areas in this region, leading to five unsuccessful wells. Through comprehensive analysis of drilled wells and seismic data, six types of horizon seismic attributes were selected. A novel approach for predicting oil-bearing zones was proposed, employing SVD–K-means nonlinear clustering for multiple seismic attribute fusion. The application results indicate: ① Singular value decomposition (SVD) technology not only reduces the correlation redundancy among seismic attribute data variables, but enables data dimensionality reduction and noise suppression, decreasing ambiguity in prediction results and enhancing reliability. ② The K-means nonlinear clustering method facilitates the nonlinear fusion of multiple seismic attribute parameters, effectively uncovering the nonlinear features of the underlying relationship between seismic attributes and reservoir oil-bearing characteristics, thereby amplifying the hydrocarbon information within the seismic attribute variables. ③ Compared to K-means, SVD–K-means demonstrates superior performance across all metrics, with an 18.4% increase in the SC coefficient, a 57.8% increase in the CH index, and a 24.7% improvement in the DB index. ④ The results of oil-bearing zone prediction using the SVD–K-means algorithm align well with the drilling outcomes in the study area and correspond to the geological patterns of hydrocarbon enrichment in this region. This has been confirmed by the high-yield industrial oil flow obtained from the newly drilled WZ6-1-A3 well. The SVD–K-means algorithm for predicting oil and gas-bearing zones provides a new approach for predicting hydrocarbon-rich areas in complex fault block structures with limited drilling and poor-quality seismic data. Full article
(This article belongs to the Section H: Geo-Energy)
Show Figures

Figure 1

Figure 1
<p>Well location map of the Weizhou Formation (W3IV) in WZ6-1 structure.</p>
Full article ">Figure 2
<p>Layer-by-layer seismic attribute plan of the Weizhou Formation (W3IV) in WZ6-1 structure.(<b>a</b>) Energy half-time; (<b>b</b>) Instantaneous phase; (<b>c</b>) Dominant frequency; (<b>d</b>) Bandwidth.</p>
Full article ">Figure 3
<p>SSE elbow method and DB index method line chart. (<b>a</b>) SSE elbow method; (<b>b</b>) DB index method.</p>
Full article ">Figure 4
<p>WZ6-1 constructs the K-means algorithm prediction diagram of the Weizhou Formation (W3IV).</p>
Full article ">Figure 5
<p>The singular value distribution curve of SVD decomposition of the Weizhou Formation (W3IV) in WZ6-1 structure is plotted.</p>
Full article ">Figure 6
<p>The SVD–K-means algorithm prediction map of the Weizhou Formation (W3IV) is constructed by WZ6-1.</p>
Full article ">
18 pages, 6578 KiB  
Article
Seismic Performance Evaluation of a Frame System Strengthened with External Self-Centering Components
by Yulin Fan, Jiaye Song, Xuelu Zhou and Hang Liu
Buildings 2024, 14(11), 3666; https://doi.org/10.3390/buildings14113666 - 18 Nov 2024
Viewed by 288
Abstract
In the context of China’s promotion of green buildings and resilient urban development, new reinforcement technologies offer significant development prospects, while traditional methods have limited effectiveness in enhancing structural resilience. To address this latter issue, this study proposes a novel reinforcement method that [...] Read more.
In the context of China’s promotion of green buildings and resilient urban development, new reinforcement technologies offer significant development prospects, while traditional methods have limited effectiveness in enhancing structural resilience. To address this latter issue, this study proposes a novel reinforcement method that involves enlarging the structural cross-section and adding external self-resetting components to improve seismic performance. While this method has been validated through quasi-static tests, limitations in terms of sample size and experimental conditions necessitate further research into the seismic performance and dynamic behavior of the reinforced framework. Consequently, this study uses finite element analysis to explore the influencing factors and dynamic characteristics of the reinforcement method. The results show that finite element modeling effectively simulates the stress characteristics of reinforced frameworks. Installing prefabricated beams significantly enhances the load-bearing capacity by 18% and reduces the residual deformation rates after earthquakes by 26%. Increased pre-tensioning of the steel strands further improves seismic resilience. This reinforcement method enables older structures lacking self-resetting capabilities to achieve some degree of self-resetting ability, and it performs well under various earthquake conditions. Full article
(This article belongs to the Section Building Structures)
Show Figures

Figure 1

Figure 1
<p>Test systems for the retrofitted concrete frames using self-centering components (the dimensions in the figure are in mm). (<b>a</b>) Photographs of the experimental site. (<b>b</b>) Section 1-1: Dimensions of the reinforced frame beams. (<b>c</b>) Section 2-2: Dimensions of the reinforced frame columns.</p>
Full article ">Figure 1 Cont.
<p>Test systems for the retrofitted concrete frames using self-centering components (the dimensions in the figure are in mm). (<b>a</b>) Photographs of the experimental site. (<b>b</b>) Section 1-1: Dimensions of the reinforced frame beams. (<b>c</b>) Section 2-2: Dimensions of the reinforced frame columns.</p>
Full article ">Figure 2
<p>Modeling concrete frame reinforcement with self-centering components in ABAQUS finite element software.</p>
Full article ">Figure 3
<p>Damage cloud map for reinforced frameworks: (<b>a</b>,<b>b</b>) cracking load and (<b>c</b>,<b>d</b>) yield load.</p>
Full article ">Figure 4
<p>Hysteresis and skeleton curves of quasi-static test and ABAQUS simulation. (<b>a</b>) Fitting of the horizontal load–displacement curve for LF1 in ABAQUS. (<b>b</b>) Fitting of the skeleton curve for LF1 in ABAQUS. (<b>c</b>) Fitting of the horizontal load–displacement curve for LF2 in ABAQUS. (<b>d</b>) Fitting of the skeleton curve for LF2 in ABAQUS. (<b>e</b>) Fitting of the horizontal load–displacement curve for LF3 in ABAQUS. (<b>f</b>) Fitting of the skeleton curve for LF3 in ABAQUS.</p>
Full article ">Figure 4 Cont.
<p>Hysteresis and skeleton curves of quasi-static test and ABAQUS simulation. (<b>a</b>) Fitting of the horizontal load–displacement curve for LF1 in ABAQUS. (<b>b</b>) Fitting of the skeleton curve for LF1 in ABAQUS. (<b>c</b>) Fitting of the horizontal load–displacement curve for LF2 in ABAQUS. (<b>d</b>) Fitting of the skeleton curve for LF2 in ABAQUS. (<b>e</b>) Fitting of the horizontal load–displacement curve for LF3 in ABAQUS. (<b>f</b>) Fitting of the skeleton curve for LF3 in ABAQUS.</p>
Full article ">Figure 5
<p>Prediction of energy dissipation in different concrete frames.</p>
Full article ">Figure 6
<p>Unit division and section setting of the OpenSees model.</p>
Full article ">Figure 7
<p>Hysteresis curve and skeleton curve of quasi-static test and OpenSees simulation. (<b>a</b>) Fitting of the horizontal load–displacement curve for LF1 in OpenSees. (<b>b</b>) Fitting of the skeleton curve for LF1 in OpenSees. (<b>c</b>) Fitting of the horizontal load–displacement curve for LF2 in OpenSees. (<b>d</b>) Fitting of the skeleton curve for LF2 in OpenSees. (<b>e</b>) Fitting of the horizontal load–displacement curve for LF3 in OpenSees. (<b>f</b>) Fitting of the skeleton curve for LF3 in OpenSees.</p>
Full article ">Figure 8
<p>OpenSees finite element full-scale model. (<b>a</b>) First-floor plan of the framework model. (<b>b</b>) Other-floor plan of the framework model. (<b>c</b>) Three-dimensional schematic diagram of frame model. (<b>d</b>) OpenSees framework model.</p>
Full article ">Figure 9
<p>Seismic waveform.</p>
Full article ">Figure 10
<p>Reaction spectrum analysis.</p>
Full article ">Figure 11
<p>Comparison of displacement time-history curves before and after reinforcement. (<b>a</b>) Frequently occurring earthquake: Chi-Chi-Taiwan. (<b>b</b>) Frequently occurring earthquake: Cape Mendocino. (<b>c</b>) Frequently occurring earthquake: artificial. (<b>d</b>) Design-basis earthquake: Chi-Chi-Taiwan. (<b>e</b>) Design-basis earthquake: Cape Mendocino. (<b>f</b>) Design-basis earthquake: artificial. (<b>g</b>) Rarely occurring earthquake: Chi-Chi-Taiwan. (<b>h</b>) Rarely occurring earthquake: Cape Mendocino. (<b>i</b>) Rarely occurring earthquake: artificial.</p>
Full article ">Figure 12
<p>Maximum displacement curve of each floor before and after structural reinforcement. (<b>a</b>) Frequently occurring earthquake: Chi-Chi-Taiwan. (<b>b</b>) Frequently occurring earthquake: Cape Mendocino. (<b>c</b>) Frequently occurring earthquake: artificial. (<b>d</b>) Design-basis earthquake: Chi-Chi-Taiwan. (<b>e</b>) Design-basis earthquake: Cape Mendocino. (<b>f</b>) Design-basis earthquake: artificial. (<b>g</b>) Rarely occurring earthquake: Chi-Chi-Taiwan. (<b>h</b>) Rarely occurring earthquake: Cape Mendocino. (<b>i</b>) Rarely occurring earthquake: artificial.</p>
Full article ">Figure 13
<p>Comparison of residual displacements before and after structural reinforcement: (<b>a</b>) Chi-Chi-Taiwan, (<b>b</b>) Cape Mendocino, and (<b>c</b>) artificial wave.</p>
Full article ">
18 pages, 7329 KiB  
Viewpoint
Study on Damage Index for Beam–Column Joints with Flush End-Plate Connections
by Jizhi Su, Jinpu Zhou, Weiran Liu, Yong Li, Haifeng Yu and Qilian Li
Buildings 2024, 14(11), 3637; https://doi.org/10.3390/buildings14113637 - 15 Nov 2024
Viewed by 309
Abstract
Experimental results from 109 flush end-plate (FEP) connections were analyzed to investigate the failure modes and damage index of FEP connections across various damage states. The present study was conducted in accordance with the performance design objectives specified in China’s GB50011-2010 Code for [...] Read more.
Experimental results from 109 flush end-plate (FEP) connections were analyzed to investigate the failure modes and damage index of FEP connections across various damage states. The present study was conducted in accordance with the performance design objectives specified in China’s GB50011-2010 Code for Seismic Design of Buildings. The observed rotation angles and corresponding rotation factors were systematically categorized using a probabilistic statistical approach, with the 95% confidence lower limit as the primary constraint. The damage states of FEP connections were classified as virtually undamaged, lightly damaged, moderately damaged, severely damaged, and joint failure. A minimum plate thickness of 12 mm and a minimum high-strength bolt diameter of 20 mm are recommended to be used for the FEP connections, in accordance with building codes in China, the United States, and Europe. Quasi-static tests of six FEP connections were conducted to validate the damage-state categorization. The results revealed that for connections undergoing moderate and severe damage, the mean rotation factor deviated from the theoretical values proposed in this study by 3.7% and 9.4%, respectively. Therefore, the damage state of FEP connections can be reliably predicted based on different rotation angles using the damage-state categorization presented herein. Full article
(This article belongs to the Section Building Structures)
Show Figures

Figure 1

Figure 1
<p>Schematic diagram of flush end-plate connection.</p>
Full article ">Figure 2
<p>EEEP method schematic diagram.</p>
Full article ">Figure 3
<p>Gap between end plate and column [<a href="#B19-buildings-14-03637" class="html-bibr">19</a>].</p>
Full article ">Figure 4
<p>Moderately damaged state. (<b>a</b>) Cracks in concrete floor [<a href="#B34-buildings-14-03637" class="html-bibr">34</a>]. (<b>b</b>) Gap between end plate and column becomes larger [<a href="#B19-buildings-14-03637" class="html-bibr">19</a>].</p>
Full article ">Figure 5
<p>Severely damaged state. (<b>a</b>) Bolt fracture [<a href="#B19-buildings-14-03637" class="html-bibr">19</a>]. (<b>b</b>) Weld cracking [<a href="#B34-buildings-14-03637" class="html-bibr">34</a>]. (<b>c</b>) Beam flange buckling [<a href="#B34-buildings-14-03637" class="html-bibr">34</a>]. (<b>d</b>) Column flange buckling [<a href="#B35-buildings-14-03637" class="html-bibr">35</a>].</p>
Full article ">Figure 6
<p>Distribution of rotation factor for severely damaged state.</p>
Full article ">Figure 7
<p>Details of specimens.</p>
Full article ">Figure 8
<p>Test setup.</p>
Full article ">Figure 9
<p>Loading system diagram.</p>
Full article ">Figure 10
<p>Failure modes of joints.</p>
Full article ">Figure 11
<p>Skeleton curves.</p>
Full article ">
Back to TopTop