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

Article Types

Countries / Regions

remove_circle_outline
remove_circle_outline
remove_circle_outline
remove_circle_outline
remove_circle_outline
remove_circle_outline

Search Results (1,589)

Search Parameters:
Keywords = Abaqus

Order results
Result details
Results per page
Select all
Export citation of selected articles as:
25 pages, 6521 KiB  
Article
Finite Element Modelling of Pressure Armour Unlocking Failure in Unbonded Flexible Risers
by Rongzhi Wei, Xiaotian Li, Murilo Augusto Vaz and Anderson Barata Custódio
J. Mar. Sci. Eng. 2025, 13(3), 411; https://doi.org/10.3390/jmse13030411 (registering DOI) - 22 Feb 2025
Viewed by 164
Abstract
Flexible pipes can be subjected to extreme bending load during installation and in situ operations, as well as high pressure from oil and gas field production fluids. Although field experience shows that the unlocking of the pressure armour layer is a rare failure [...] Read more.
Flexible pipes can be subjected to extreme bending load during installation and in situ operations, as well as high pressure from oil and gas field production fluids. Although field experience shows that the unlocking of the pressure armour layer is a rare failure mode, its consequences are potentially catastrophic. To investigate the unlocking mechanism of the pressure armour layer, a 3D nonlinear finite element model is developed, which contains three layers: a pressure armour layer with a “Zeta” shape cross-section, a polymer layer, and an analytical cylindrical surface representing the radial support from the carcass. The analysis is carried out with the ABAQUS Dynamic Explicit solver using a semi-automatic mass scaling factor. The study focuses on the possibility of pressure armour unlocking due to bending load, where the effects of geometric features of the pressure armour layer and internal pressure value on unlocking are investigated by varying corresponding parameters. The influence of these variables are discussed at the end. Ultimately, in conjunction with the numerical analysis results, critical reflections on the shortcomings of the model and suggestions for improvement are presented based on the numerical analysis results. Full article
(This article belongs to the Special Issue Advanced Research in Flexible Riser and Pipelines)
23 pages, 8778 KiB  
Article
A Novel Approach to the Design of Distributed Dynamic Vibration Absorbers for Plates Subjected to Classical and Elastic Edge Conditions
by Yuan Du, Yuhang Tang, Chenyu Fan, Yucheng Zou, Zhen Bao and Yong Ma
J. Mar. Sci. Eng. 2025, 13(3), 401; https://doi.org/10.3390/jmse13030401 - 21 Feb 2025
Viewed by 101
Abstract
Plate structures are the main components of offshore platforms and ships in engineering applications. The vibration control of the low-frequency mode of plate structures has always been a meaningful research object in marine science and engineering. Due to their low cost and good [...] Read more.
Plate structures are the main components of offshore platforms and ships in engineering applications. The vibration control of the low-frequency mode of plate structures has always been a meaningful research object in marine science and engineering. Due to their low cost and good performance, dynamic vibration absorbers are widely used. To enhance the design efficiency of dynamic vibration absorbers, a mathematical model was developed for plates with dynamic vibration absorbers under different boundary constraints. To overcome the discontinuity of the displacement function, auxiliary series were introduced. In addition, the efficiency of resolving the plate structure’s equivalent mass was significantly improved compared with when using FEM software Abaqus 6.14. The validity of the proposed mathematical model was confirmed in comparison with related studies, the FEM results, and the experimental results. Considering the mathematical model and design approach proposed in the current paper, more research on the vibration control of plates subjected to clamped and elastic boundary conditions should be performed. The mathematical model and findings in the design process may have positive implications for the control of the vibration of plate structures in marine science and engineering. Full article
(This article belongs to the Section Ocean Engineering)
Show Figures

Figure 1

Figure 1
<p>Ship plate structure.</p>
Full article ">Figure 2
<p>The schematic diagram of rectangular plate with a dynamic vibration absorber.</p>
Full article ">Figure 3
<p>An illustration of the possible discontinuities of the conventional Fourier series at the endpoints.</p>
Full article ">Figure 4
<p>An illustration of the principle of how to eliminate the possible discontinuities of the improved Fourier series at the endpoints.</p>
Full article ">Figure 5
<p>Equivalent mass solution procedure of a plate structure.</p>
Full article ">Figure 6
<p>The first natural frequency of a four-sided clamped rectangular plate (<span class="html-italic">f</span> = 8.97 Hz).</p>
Full article ">Figure 7
<p>The variation tendency of equivalent mass with additional mass at point A.</p>
Full article ">Figure 8
<p>The relationship between the dimensionless frequency parameters of the plate and the spring stiffness.</p>
Full article ">Figure 9
<p>Experimental devices and test model.</p>
Full article ">Figure 10
<p>Mode shapes of the cantilever plate (left side: test results; right side: current method).</p>
Full article ">Figure 11
<p>Schematic diagram of multi-point excitation and assessment points.</p>
Full article ">Figure 12
<p>Vibration response of the plate for different frequency ratios of the dynamic vibration absorber.</p>
Full article ">Figure 13
<p>Vibration response of the plate when the damping of dynamic vibration absorber is optimized.</p>
Full article ">Figure 14
<p>Vibration response of the plate for different installation positions of the dynamic vibration absorber.</p>
Full article ">Figure 15
<p>The design procedure of the dynamic vibration absorber for the plate.</p>
Full article ">Figure 16
<p>Schematic diagram of multi-point excitation points and assessment points.</p>
Full article ">Figure 17
<p>Frequency response curve under multi-point excitation load.</p>
Full article ">Figure 18
<p>Mounting points of distributed dynamic vibration absorber.</p>
Full article ">Figure 19
<p>Comparison of the vibration response curves before and after installation of the distributed dynamic vibration absorber.</p>
Full article ">Figure 19 Cont.
<p>Comparison of the vibration response curves before and after installation of the distributed dynamic vibration absorber.</p>
Full article ">Figure 20
<p>Frequency response curves under multi-point excitation load for classical and elastic boundary condition.</p>
Full article ">Figure 21
<p>Mounting points of distributed dynamic vibration absorbers under elastic boundary conditions.</p>
Full article ">Figure 22
<p>Effect of the dynamic vibration absorber under elastic boundary conditions.</p>
Full article ">Figure 22 Cont.
<p>Effect of the dynamic vibration absorber under elastic boundary conditions.</p>
Full article ">
21 pages, 11401 KiB  
Article
Numerical Analysis of the Cyclic Behavior of Reinforced Concrete Columns Incorporating Rubber
by Mohammed A. M. Ahmed, Heba A. Mohamed, Hilal Hassan, Ayman El-Zohairy and Mohamed Emara
J. Compos. Sci. 2025, 9(3), 95; https://doi.org/10.3390/jcs9030095 - 21 Feb 2025
Viewed by 140
Abstract
A numerical analysis of rubberized reinforced concrete columns’ performance under cyclic loading is presented in this study. Three different concrete blends (M1, M2, and M3) were chosen based on the volume of fine aggregate replaced by varying percentages of crumb rubber (CR) (0%, [...] Read more.
A numerical analysis of rubberized reinforced concrete columns’ performance under cyclic loading is presented in this study. Three different concrete blends (M1, M2, and M3) were chosen based on the volume of fine aggregate replaced by varying percentages of crumb rubber (CR) (0%, 10%, and 15%). Under cyclic loads, three groups of rubberized reinforced concrete (RRC) columns with circular, square, and rectangular cross-sections and heights of 1.5 m and 2.0 m were analyzed using the finite element software ABAQUS. The proposed model effectively predicts the behavior of rubberized reinforced concrete columns under cyclic loading. Additionally, these columns demonstrate improved performance in lateral displacement, displacement ductility, and damping ratio, with only a slight reduction in lateral load capacity. For the circular columns with a height of 1.5 m, the displacement ductility increased by 47.8% and 89.0% when the fine aggregates were replaced with 10% and 15% CR, respectively. Similarly, for square columns of the same height, the displacement ductility increased by 18.7% and 26.7% with 10% and 15% CR, respectively. The rectangular specimens exhibited enhancements of 34.74% and 58.95%, respectively. Although the analyzed rubberized reinforced concrete columns experienced slight reductions in the lateral load capacity compared to the non-CR columns, the cyclic damage resistance was notably improved. Full article
(This article belongs to the Special Issue Theoretical and Computational Investigation on Composite Materials)
Show Figures

Figure 1

Figure 1
<p>Details of the circular columns of Group 1.</p>
Full article ">Figure 2
<p>Details of the square columns of Group 2.</p>
Full article ">Figure 3
<p>Details of the rectangular columns of Group 3.</p>
Full article ">Figure 4
<p>Crumb rubber particles [<a href="#B18-jcs-09-00095" class="html-bibr">18</a>].</p>
Full article ">Figure 5
<p>Effects of CR on the mechanical properties of concrete.</p>
Full article ">Figure 6
<p>Concrete damage plasticity model provided by ABAQUS [<a href="#B18-jcs-09-00095" class="html-bibr">18</a>].</p>
Full article ">Figure 7
<p>FE simulation utilizes constitutive models of materials [<a href="#B18-jcs-09-00095" class="html-bibr">18</a>].</p>
Full article ">Figure 8
<p>The loading protocol [<a href="#B22-jcs-09-00095" class="html-bibr">22</a>].</p>
Full article ">Figure 9
<p>Loading and boundary conditions of the FE models. These arrows represent the directions of the axial load and lateral displacement, which are provided on the figure.</p>
Full article ">Figure 10
<p>RC column details [<a href="#B22-jcs-09-00095" class="html-bibr">22</a>].</p>
Full article ">Figure 11
<p>Results of experimental and FEM RC column.</p>
Full article ">Figure 12
<p>RRC column details.</p>
Full article ">Figure 13
<p>Results of the experimental and FE results for the analyzed RRC column.</p>
Full article ">Figure 14
<p>Hysteretic behavior of circular columns (Group 1).</p>
Full article ">Figure 15
<p>Hysteretic behavior of square columns (Group 2).</p>
Full article ">Figure 16
<p>Hysteretic behavior of rectangular columns (Group 3).</p>
Full article ">Figure 17
<p>Backbone curve for columns.</p>
Full article ">Figure 18
<p>Equivalent viscous damping and displacement ductility of columns.</p>
Full article ">Figure 19
<p>Effect of crumb rubber on equivalent viscous damping and displacement ductility of columns.</p>
Full article ">
16 pages, 15805 KiB  
Article
Study on Damage Mechanism of Fiber Concrete with Initial Pores
by Ankui Hu, Xinyu Du, Fei Wang, Junjie Li, Tianlong Zhang and Yajing Li
Materials 2025, 18(5), 916; https://doi.org/10.3390/ma18050916 - 20 Feb 2025
Viewed by 137
Abstract
Currently, fiber-reinforced concrete, as a building material, is widely used in highway bridges and tunnel linings, and it has become a global research hotspot, with indoor tests, numerical simulations, performance studies, and application scenarios surrounding it. Many researchers have conducted experiments and analyses [...] Read more.
Currently, fiber-reinforced concrete, as a building material, is widely used in highway bridges and tunnel linings, and it has become a global research hotspot, with indoor tests, numerical simulations, performance studies, and application scenarios surrounding it. Many researchers have conducted experiments and analyses on the damage patterns of fiber-reinforced concrete under different conditions. However, there is relatively little research on the mechanical properties of fiber-reinforced concrete that already contains initial damage. This article establishes a micro-model composed of aggregates, mortar, and interface layers using MATLAB. It introduces the CDP (Concrete Damage Plasticity) constitutive equation for fiber-reinforced concrete and uses the least squares method to fit and validate the equation. After model validation, uniaxial compression tests are conducted on models with different initial porosities using the ABAQUS (2023) software, resulting in changes in crack damage, peak stress, and elastic modulus mechanical properties. The conclusions are as follows: The improved characteristic structure curve using the least squares method fits the experimental results well, and the rationality of the algorithm was verified by comparing it with physical tests. As the porosity increased from 2% to 8%, the peak stress decreased from 98.6% to 70.5% compared to non-porous fiber concrete with a significant rate of decrease of about 30%. After considering the strain rate, the peak stress increased slowly with increasing strain rate, but the elastic modulus increased at a significant rate, with a 1.26 times higher elastic modulus at a strain rate of 100 than at a strain rate of 10−2. This result provides a certain theoretical basis for the mechanical properties and damage modes of fiber-containing concrete in practical engineering. Full article
(This article belongs to the Section Construction and Building Materials)
Show Figures

Figure 1

Figure 1
<p>Flowchart for random aggregate and pore generation.</p>
Full article ">Figure 2
<p>Parameters of the components of the microstructure of concrete.</p>
Full article ">Figure 3
<p>Comparison chart of constitutive curves and experimental curves.</p>
Full article ">Figure 4
<p>Comparison chart of the fitted curve and the experimental curve.</p>
Full article ">Figure 5
<p>Comparison of uniaxial compression stress–strain curves with different porosities.</p>
Full article ">Figure 6
<p>Comparison of displacements at different porosities.</p>
Full article ">Figure 7
<p>Crack damage evolution diagram for concrete with different pore fibers.</p>
Full article ">Figure 8
<p>Graph of the relationship between the peak stress of fiber-reinforced concrete with initial porosity and strain rate.</p>
Full article ">Figure 9
<p>Graph of the relationship of the elastic modulus of fiber-reinforced concrete with initial porosity, porosity, and strain rate.</p>
Full article ">
27 pages, 6813 KiB  
Article
Application of Unprocessed Waste Tyres in Pavement Base Structures: A Study on Deformation and Stress Analysis Using Finite Element Simulation
by Baoying Shen, Hui Tian, Wenruo Fan, Lu Zhang and Hui Wang
Materials 2025, 18(4), 914; https://doi.org/10.3390/ma18040914 - 19 Feb 2025
Viewed by 223
Abstract
In this study, numerical simulations using the Abaqus finite element model were performed to evaluate the effects of incorporating waste tyres of varying sizes into the base layer as part of a coupled tyre–pavement structure. The tyre-reinforced structure demonstrated superior deformation resilience, attributed [...] Read more.
In this study, numerical simulations using the Abaqus finite element model were performed to evaluate the effects of incorporating waste tyres of varying sizes into the base layer as part of a coupled tyre–pavement structure. The tyre-reinforced structure demonstrated superior deformation resilience, attributed to the hyperelastic properties of tyre rubber, underscoring its potential for applications where deformation recovery is essential. For achieving a uniform settlement, the entire tyre stacking scheme is recommended, whereas the one-third tyre configuration is ideal for minimising displacement. The one-half tyre configuration provides a balanced approach, optimising resource utilisation for structures with moderate performance requirements. The inclusion of tyres increases the equivalent stress within the cement-stabilised gravel layer beneath the tyre, and this effect is less pronounced with smaller tyre sizes. Notably, the projected portion of the tyre tread enhances the bearing capacity of the base structure, improving the load distribution and overall structural performance. The middle and bottom surface layers were identified as the most critical for controlling deformation and stress distribution, while a moderate modulus is advised for the surface course to achieve a balance between deformation control and stress uniformity. The integration of high-modulus layers with tyre reinforcement offers an optimised solution for both deformation management and stress distribution. This study highlights the potential of tyre-reinforced pavements as an innovative and sustainable construction practice, particularly suited for light to moderate traffic conditions. Further research is recommended to explore the long-term environmental and economic benefits, as well as the impacts of tyre composition and ageing on performance. Full article
Show Figures

Figure 1

Figure 1
<p>A 3D model of the pavement structure.</p>
Full article ">Figure 2
<p>Tyre modelling. (<b>a</b>) Planar parameters; (<b>b</b>) 3D modelling.</p>
Full article ">Figure 3
<p>Tyre arrangement method. (<b>a</b>) Whole; (<b>b</b>) half; (<b>c</b>) one-third (third); (<b>d</b>) one-quarter (quarter).</p>
Full article ">Figure 4
<p>Conversion of tyre–road contact shape and equivalent area.</p>
Full article ">Figure 5
<p>Grid division schematic diagram. (<b>a</b>) Road grid division; (<b>b</b>) tyre grid division.</p>
Full article ">Figure 6
<p>Typical pavement structure deformation maps. (<b>a</b>) Original road; (<b>b</b>) one whole tyre placed; (<b>c</b>) one 1/4 tyre placed; (<b>d</b>) three whole tyres placed.</p>
Full article ">Figure 7
<p>Displacement results of asphalt mixture layer bottom. (<b>a</b>) One tyre placed; (<b>b</b>) three tyres placed.</p>
Full article ">Figure 7 Cont.
<p>Displacement results of asphalt mixture layer bottom. (<b>a</b>) One tyre placed; (<b>b</b>) three tyres placed.</p>
Full article ">Figure 8
<p>Stress cloud of some pavement structures. (<b>a</b>) Only road (OnlyR); (<b>b</b>) one whole tyre placed; (<b>c</b>) one 1/4 tyre placed; (<b>d</b>) three whole tyres placed.</p>
Full article ">Figure 8 Cont.
<p>Stress cloud of some pavement structures. (<b>a</b>) Only road (OnlyR); (<b>b</b>) one whole tyre placed; (<b>c</b>) one 1/4 tyre placed; (<b>d</b>) three whole tyres placed.</p>
Full article ">Figure 9
<p>Equivalent force variation curves in the depth direction. (<b>a</b>) One tyre placed; (<b>b</b>) three tyres placed. Note: arrow positions represent the location of structural layer changes.</p>
Full article ">Figure 9 Cont.
<p>Equivalent force variation curves in the depth direction. (<b>a</b>) One tyre placed; (<b>b</b>) three tyres placed. Note: arrow positions represent the location of structural layer changes.</p>
Full article ">Figure 10
<p>Equivalent stress distribution results for Static Loads. (<b>a</b>) Top; (<b>b</b>) middle; (<b>c</b>) bottom.</p>
Full article ">Figure 10 Cont.
<p>Equivalent stress distribution results for Static Loads. (<b>a</b>) Top; (<b>b</b>) middle; (<b>c</b>) bottom.</p>
Full article ">Figure 11
<p>Schematic of data observation positions.</p>
Full article ">Figure 12
<p>Displacement results of the top surface of the road.</p>
Full article ">Figure 13
<p>Equivalent stress distribution results for Moving Loads. (<b>a</b>) Top; (<b>b</b>) middle; (<b>c</b>) bottom.</p>
Full article ">Figure 13 Cont.
<p>Equivalent stress distribution results for Moving Loads. (<b>a</b>) Top; (<b>b</b>) middle; (<b>c</b>) bottom.</p>
Full article ">Figure 14
<p>Impact of factors on MDBSL.</p>
Full article ">Figure 15
<p>Impact of factors on EDESDD.</p>
Full article ">
22 pages, 9353 KiB  
Article
Numerical Investigation of the Axial Compressive Behavior of a Novel L-Shaped Concrete-Filled Steel Tube Column
by Fujian Yang, Yi Bao, Muzi Du and Xiaoshuang Li
Materials 2025, 18(4), 897; https://doi.org/10.3390/ma18040897 - 19 Feb 2025
Viewed by 155
Abstract
A novel L-shaped concrete-filled steel tube (CFST) column is proposed in this study. A finite element model of the column is developed using ABAQUS software to analyze its load transfer mechanism and axial compressive behavior. The effects of factors such as the steel [...] Read more.
A novel L-shaped concrete-filled steel tube (CFST) column is proposed in this study. A finite element model of the column is developed using ABAQUS software to analyze its load transfer mechanism and axial compressive behavior. The effects of factors such as the steel strength, steel tube thickness, support plate configuration, and perforation of the support plates on the compressive performance of the column are investigated. The simulation results reveal that the column exhibits robust axial compressive performance. Increasing the steel strength and incorporating support plates (SP) effectively enhance the column’s compressive bearing capacity and positively influence the bearing capacity coefficient (δ). However, increasing the steel tube thickness results in a reduction in δ, indicating that the rate of increase in the bearing capacity diminishes with increasing thickness. The failure mode is primarily characterized by local buckling in the midsection of the steel tube’s concave corner. Measures such as increasing the steel strength and tube thickness and the use of support plates help to mitigate buckling at the concave corner, improve concrete confinement, and enhance the overall compressive performance of the column. Full article
(This article belongs to the Section Construction and Building Materials)
Show Figures

Figure 1

Figure 1
<p>Cross-sectional forms of traditional CFST columns.</p>
Full article ">Figure 2
<p>Schematic of steel tube column.</p>
Full article ">Figure 3
<p>Dimensional design of L-shaped steel tube column.</p>
Full article ">Figure 4
<p>Constitutive models for the CDP: (<b>a</b>) compressive; (<b>b</b>) tensile.</p>
Full article ">Figure 5
<p>Typical yield surfaces for the CDP: (<b>a</b>) yield surfaces in the deviatoric plane; (<b>b</b>) yield surface in plane stress.</p>
Full article ">Figure 6
<p>Stress–strain curve of steel.</p>
Full article ">Figure 7
<p>Finite element model of irregular L-shaped CFST column.</p>
Full article ">Figure 8
<p>Load–displacement relationship comparison for cross-shaped CFST column: (<b>a</b>) load–vertical displacement relationship curve of CFST column; (<b>b</b>) load–horizontal displacement relationship curve at midsection of CFST column.</p>
Full article ">Figure 9
<p>Load–displacement relationships for L-shaped CFST columns with different structural forms.</p>
Full article ">Figure 10
<p>Stress distribution contour map of the steel tube at 85% of the peak load: (<b>a</b>) CFST-M-2; (<b>b</b>) CFST-S-2; (<b>c</b>) CFST-S-3.</p>
Full article ">Figure 11
<p>Stress distribution of steel tube at maximum displacement: (<b>a</b>) CFST-M-2; (<b>b</b>) CFST-S-2; (<b>c</b>) CFST-S-3.</p>
Full article ">Figure 12
<p>Location of unidirectional eccentric load.</p>
Full article ">Figure 13
<p>Comparison of load–displacement curves under eccentric and axial compression conditions: (<b>a</b>) CFST-M-2; (<b>b</b>) CFST-S-2; (<b>c</b>) CFST-S-3.</p>
Full article ">Figure 14
<p>Stress distribution of steel tube at column end at maximum displacement under eccentric load: (<b>a</b>) CFST-M-2; (<b>b</b>) CFST-S-2; (<b>c</b>) CFST-S-3.</p>
Full article ">Figure 15
<p>Distribution of equivalent plastic strain in the concrete column at the maximum displacement at the column base under eccentric loading: (<b>a</b>) CFST-M-2; (<b>b</b>) CFST-S-2; (<b>c</b>) CFST-S-3.</p>
Full article ">Figure 16
<p>Load–displacement relationship in L-shaped CFST columns with different steel strengths.</p>
Full article ">Figure 17
<p>Stress distribution of steel tube at column end at maximum displacement under eccentric load: (<b>a</b>) CFST-M-1; (<b>b</b>) CFST-M-2; (<b>c</b>) CFST-M-3.</p>
Full article ">Figure 18
<p>Load–displacement relationship for L-shaped CFST columns with different steel tube thicknesses.</p>
Full article ">Figure 19
<p>Stress distribution of steel tube at maximum displacement: (<b>a</b>) CFST-T-1; (<b>b</b>) CFST-M-2; (<b>c</b>) CFST-T-3.</p>
Full article ">Figure 20
<p>Load–displacement relationship for L-shaped CFST columns with different numbers of supporting plates.</p>
Full article ">Figure 21
<p>Stress distribution of steel tube at maximum displacement: (<b>a</b>) CFST-S-1; (<b>b</b>) CFST-S-2.</p>
Full article ">Figure 22
<p>Stress distribution of the supporting plate: (<b>a</b>) CFST-S-2; (<b>b</b>) CFST-S-3.</p>
Full article ">Figure 23
<p>Comparison of load–displacement relationships for L-shaped CFST columns with different column heights.</p>
Full article ">Figure 24
<p>Equivalent plastic strain (PEEQ) distribution contour for L-shaped concrete columns: (<b>a</b>) CFST-M-2; (<b>b</b>) CFST-H-1.</p>
Full article ">Figure 25
<p>Concrete compression damage distribution: (<b>a</b>) CFST-M-1; (<b>b</b>) CFST-M-2; (<b>c</b>) CFST-M-3; (<b>d</b>) CFST-T-1; (<b>e</b>) CFST-T-3; (<b>f</b>) CFST-S-1; (<b>g</b>) CFST-S-2; (<b>h</b>) CFST-S-3; (<b>i</b>) CFST-H-1.</p>
Full article ">
18 pages, 7632 KiB  
Article
Research on the Fine Control of the Influence of Pipe-Jacking Parameter Deviation on Surrounding Stratum Deformation
by Tianlong Zhang, Guoqing Chen, Ping Lu and Dongqing Nie
Appl. Sci. 2025, 15(4), 2208; https://doi.org/10.3390/app15042208 - 19 Feb 2025
Viewed by 199
Abstract
Based on the Zhuyuan–Bailonggang sewage interconnection pipe project in Shanghai, the ABAQUS finite element software was used in numerical simulations to study the fine control of stratum disturbances caused by pipe jacking parameter deviation in soft soil areas. Combining the simulation results with [...] Read more.
Based on the Zhuyuan–Bailonggang sewage interconnection pipe project in Shanghai, the ABAQUS finite element software was used in numerical simulations to study the fine control of stratum disturbances caused by pipe jacking parameter deviation in soft soil areas. Combining the simulation results with onsite measured data, the Peck formula was used to predict surface settlement. The results indicate the following: (1) The jacking speed and face pressure are negatively correlated with surface settlement. Under the maximum positive deviation and negative deviations in the jacking speed, after the tail passes through the monitoring section D0 16 ring, the maximum value of settlement at point B8 increases by 21.6% and decreases by 12.8%, respectively. Increasing the jacking speed increases the area with stress change ratio R < 0 at monitoring section D0, and the arch foot at the tail of the pipe jacking machine decreases the surface settlement. In contrast, when the face pressure deviates from its average value, the variation range is less than 1%. (2) The pipe slurry coefficient and surface subsidence are positively correlated. Under the maximum positive deviation and the maximum negative deviation, the tail passes through the monitoring section D0 16 ring, and the maximum settlement value at B8 decreases by 4.9% and increases by 16.5%, respectively. The increase in the coefficient reduces the area with R < 0 at D0 and increases the surface settlement. (3) In the order of descending strength, surface settlement is affected by the jacking speed, slurry friction coefficient, and face pressure. (4) To predict the maximum surface settlement value due to deviations in the jacking parameters, the Peck formula was modified using a correction factor α ranging from 0.6 to 3.0 and a settlement trough width correction factor β ranging from 1.6 to 4.0. The modified prediction curve is in closer agreement with the actual conditions. Full article
Show Figures

Figure 1

Figure 1
<p>Plane layout of monitoring points.</p>
Full article ">Figure 2
<p>Surface subsidence monitoring curves: (<b>a</b>) lateral settlement; (<b>b</b>) longitudinal settlement.</p>
Full article ">Figure 3
<p>Jacking parameter fluctuation diagram.</p>
Full article ">Figure 4
<p>Finite element model and size.</p>
Full article ">Figure 5
<p>Calculated curve and measured results of ground settlement at D<sub>0</sub>.</p>
Full article ">Figure 6
<p>Influence of the jacking speed deviation on the ground settlement.</p>
Full article ">Figure 7
<p>Change in the transverse settlement tank under different jacking speeds.</p>
Full article ">Figure 8
<p>Comparison diagram of the vertical stress on the soil arch under two kinds of jacking speeds: (<b>a</b>) <math display="inline"><semantics> <mi>v</mi> </semantics></math> = 2 ring/d; (<b>b</b>) <math display="inline"><semantics> <mi>v</mi> </semantics></math> = 10 ring/d.</p>
Full article ">Figure 9
<p>Influence of tunnel face pressure deviation on ground settlement.</p>
Full article ">Figure 10
<p>Comparison diagram of the vertical stress on the soil arch under two kinds of tunnel face pressures: (<b>a</b>) <math display="inline"><semantics> <mrow> <msub> <mi>p</mi> <mi>N</mi> </msub> <mo>=</mo> <mn>100</mn> <mi>kPa</mi> </mrow> </semantics></math>; (<b>b</b>) <math display="inline"><semantics> <mrow> <msub> <mi>p</mi> <mi>N</mi> </msub> <mo>=</mo> <mn>140</mn> <mi>kPa</mi> </mrow> </semantics></math>.</p>
Full article ">Figure 11
<p>Influence of the friction coefficient deviation on ground settlement.</p>
Full article ">Figure 12
<p>Change diagram of the transverse settlement tank under different coefficients of friction.</p>
Full article ">Figure 13
<p>Comparison diagram of the vertical stress on the soil arch under two coefficients of friction: (<b>a</b>) <math display="inline"><semantics> <mrow> <mi>μ</mi> <mo>=</mo> <mn>0.01</mn> </mrow> </semantics></math>; (<b>b</b>) <math display="inline"><semantics> <mrow> <mi>μ</mi> <mo>=</mo> <mn>0.09</mn> </mrow> </semantics></math>.</p>
Full article ">Figure 14
<p>Normalized vertical surface displacement caused by parameter deviation.</p>
Full article ">Figure 15
<p>Peck fitting, prediction, and comparison of the measured data curves.</p>
Full article ">Figure 16
<p>Peck prediction correction, fitting, and comparison of measured data curves.</p>
Full article ">Figure 17
<p>Distribution intervals of the α and β data: (<b>a</b>) <math display="inline"><semantics> <mi>α</mi> </semantics></math> data distribution diagram; (<b>b</b>) <math display="inline"><semantics> <mi>β</mi> </semantics></math> data distribution diagram.</p>
Full article ">Figure 18
<p>Comparison of the corrected Peck upper and lower limit curves with the measured data.</p>
Full article ">
29 pages, 34281 KiB  
Article
Bio-Inspired Thin-Walled Straight and Tapered Tubes with Variable Designs Subjected to Multiple Impact Angles for Building Constructions
by Quanjin Ma, Nor Hazwani Mohd Yusof, Santosh Kumar Sahu, Yiheng Song, Nabilah Afiqah Mohd Radzuan, Bo Sun, Ahmad Yunus Nasution, Alagesan Praveen Kumar and Mohd Ruzaimi Mat Rejab
Buildings 2025, 15(4), 620; https://doi.org/10.3390/buildings15040620 - 17 Feb 2025
Viewed by 213
Abstract
Thin-walled structures are extensively utilized in construction because of their lightweight nature and excellent energy absorption efficiency, especially under dynamic loads. Improving the energy-absorbing performance of thin-walled structures by inspiring natural multi-cell designs is a sufficient approach. This paper investigates the energy-absorbing characteristics [...] Read more.
Thin-walled structures are extensively utilized in construction because of their lightweight nature and excellent energy absorption efficiency, especially under dynamic loads. Improving the energy-absorbing performance of thin-walled structures by inspiring natural multi-cell designs is a sufficient approach. This paper investigates the energy-absorbing characteristics of variable novel cross-section designs of thin-walled structures subjected to oblique impact loading. Straight and tapered types with seven cross-sectional designs of novel thin-walled structures were studied. The nonlinear ABAQUS/Explicit software 6.13 version was implemented to analyze the crashworthiness behaviors for the proposed variable cross-section designs under different loading angles. The crushing behaviors of the proposed thin-walled structures were examined for various wall thicknesses of 0.5 mm, 1.5 mm, and 2.5 mm and impact loading angles of 0°, 15°, 30°, and 45°. It was determined that the energy-absorbing characteristics of novel thin-walled structures can be efficiently controlled by varying two geometries and seven cross-section designs. A multi-criteria decision-making method (MCDM) using a complex proportional assessment method (COPRAS) was performed to select the optimum thin-walled structures with cross-section designs. It was shown that a tapered square thin-walled structure with 2.5 mm thickness had the best crashworthiness performances with energy absorption (EA) of 11.01 kJ and specific energy absorption (SEA) of 20.32 kJ/kg under a 30° impact angle. Moreover, the results indicated that the EA of the thin-walled structure decreased with the increase in the impact loading angle. In addition, with the increase in the impact loading angle, the peak crushing force (PCF) decreased and reflected the reduction in energy absorbed at a larger angle. The MCDM method in conjunction with the COPRAS method is proposed; it provides valuable insights for safer and more resilient building construction. Full article
(This article belongs to the Special Issue Bionic Materials and Structures in Civil Engineering)
Show Figures

Figure 1

Figure 1
<p>Thin-walled structure as tubular roof truss element in civil engineering under multi-angle impacts.</p>
Full article ">Figure 2
<p>Bio-inspired design concepts from the cactus for building constructions in civil engineering.</p>
Full article ">Figure 3
<p>Seven designs of the thin-walled structures were used in this study: (<b>a</b>) straight tube; (<b>b</b>) tapered tube.</p>
Full article ">Figure 3 Cont.
<p>Seven designs of the thin-walled structures were used in this study: (<b>a</b>) straight tube; (<b>b</b>) tapered tube.</p>
Full article ">Figure 4
<p>Procedure of finite element modeling: (<b>a</b>) Example of impact angle and boundary conditions of tapered nonagon tubes; (<b>b</b>) mesh convergence sensitivity analysis of thin-walled tube in terms of CPU time and percentage of correlation; (<b>c</b>) oblique impact angles from thin-walled truss in civil engineering.</p>
Full article ">Figure 5
<p>Deformation modes of thin-walled straight tubes with 0.5 mm wall thickness under 0°, 15°, 30°, and 45° impact angles.</p>
Full article ">Figure 6
<p>Deformation modes of thin-walled straight tubes with 1.5 mm wall thickness under 0°, 15°, 30°, and 45° impact angles.</p>
Full article ">Figure 7
<p>Deformation modes of thin-walled straight tubes with 2.5 mm wall thickness under 0°, 15°, 30°, and 45° impact angles.</p>
Full article ">Figure 8
<p>Deformation modes of thin-walled tapered tubes with 0.5 mm wall thickness under 0°, 15°, 30°, and 45° impact angles.</p>
Full article ">Figure 9
<p>Deformation modes of thin-walled tapered tubes with 1.5 mm wall thickness under 0°, 15°, 30°, and 45° impact angles.</p>
Full article ">Figure 10
<p>Deformation modes of thin-walled tapered tubes with 2.5 mm wall thickness under 0°, 15°, 30°, and 45° impact angles.</p>
Full article ">Figure 11
<p>Load–displacement curves of thin-walled straight tubes with 0.5 mm wall thickness subjected to multiple loading angles: (<b>a</b>) 0°; (<b>b</b>) 15°; (<b>c</b>) 30°; (<b>d</b>) 45°.</p>
Full article ">Figure 12
<p>Load–displacement curves of thin-walled straight tubes with 1.5 mm wall thickness subjected to multiple loading angles: (<b>a</b>) 0°; (<b>b</b>) 15°; (<b>c</b>) 30°; (<b>d</b>) 45°.</p>
Full article ">Figure 13
<p>Load–displacement curves of thin-walled straight tubes with 2.5 mm wall thickness subjected to multiple loading angles: (<b>a</b>) 0°; (<b>b</b>) 15°; (<b>c</b>) 30°; (<b>d</b>) 45°.</p>
Full article ">Figure 14
<p>Effect of impact angle and wall thickness of specific energy absorption (SEA) on thin-walled straight tubes under three wall thicknesses: (<b>a</b>) 0.5 mm; (<b>b</b>) 1.5 mm; (<b>c</b>) 2.5 mm.</p>
Full article ">Figure 15
<p>Effect of impact angle and wall thickness of specific energy absorption (SEA) on thin-walled tapered tubes under three wall thicknesses: (<b>a</b>) 0.5 mm; (<b>b</b>) 1.5 mm; (<b>c</b>) 2.5 mm.</p>
Full article ">Figure 16
<p>Results of SEA of thin-walled tubes as a function of impact angle and wall thickness: (<b>a</b>) straight type; (<b>b</b>) tapered type.</p>
Full article ">Figure 17
<p>Result of overall SEAα with thin-walled straight and tapered tubes with seven geometry profiles: (<b>a</b>) CASE I; (<b>b</b>) CASE II.</p>
Full article ">Figure 18
<p>Effect of impact angle and wall thickness of PCF on thin-walled straight tubes with three wall thicknesses: (<b>a</b>) 0.5 mm; (<b>b</b>) 1.5 mm; (<b>c</b>) 2.5 mm.</p>
Full article ">Figure 19
<p>Effect of impact angle and wall thickness of PCF on thin-walled tapered tubes with three wall thicknesses: (<b>a</b>) 0.5 mm; (<b>b</b>) 1.5 mm; (<b>c</b>) 2.5 mm.</p>
Full article ">Figure 20
<p>Results of COPRAS method for MCDM process on straight and tapered tubes with different designs: (<b>a</b>) optimum design; (<b>b</b>) worst design.</p>
Full article ">Figure 21
<p>Thin-walled structures used in civil engineering applications: (<b>a</b>) truss bridge; (<b>b</b>) roof truss structural framework; (<b>c</b>) transmission tower; (<b>d</b>) “Eye of Shenzhen” of Gangxia North Hub Station.</p>
Full article ">
21 pages, 33789 KiB  
Article
Numerical Simulation of the Gas Flow of Combustion Products from Ignition in a Solid Rocket Motor Under Conditions of Propellant Creep
by Yin Zhang, Zhensheng Sun, Yu Hu, Yujie Zhu, Xuefeng Xia, Huang Qu and Bo Tian
Aerospace 2025, 12(2), 153; https://doi.org/10.3390/aerospace12020153 - 17 Feb 2025
Viewed by 211
Abstract
The development of modern solid rocket technology with high-performance and high-loading ratio propellants places higher requirements on the safety and stability of the solid rocket motor. The propellant of the solid rocket motor will creep during long-term vertical storage, which may adversely affect [...] Read more.
The development of modern solid rocket technology with high-performance and high-loading ratio propellants places higher requirements on the safety and stability of the solid rocket motor. The propellant of the solid rocket motor will creep during long-term vertical storage, which may adversely affect its regular operation. The ignition transient process is a critical phase in the operation of solid rocket motors. The Abaqus v.2022 finite element simulation software is used to analyze the ignition transient under propellant creep conditions and obtain the deformed combustion chamber profile. Then, we use a high-precision finite volume solver developed independently to simulate the flow field during the ignition process. In the simulation, we adopt the surface temperature of the propellant column reaching the ignition threshold as the ignition criterion, considering the heat transfer process of the propellant column instead of using the near-wall gas temperature to obtain the set temperature. Simulation results under different creep conditions reveal that the deformation of the propellant grains progressively intensifies as the solid rocket motor’s storage duration increases. This leads to a delayed initial ignition time of the propellant, an advancement of the overall ignition transient process, and an increased pressurization rate during ignition, which can affect the structure and regular operation of the motor. The research results provide design guidance and theoretical support for the design and life prediction of solid rocket motors. Full article
(This article belongs to the Section Astronautics & Space Science)
Show Figures

Figure 1

Figure 1
<p>Solid rocket motor propellant grain model.</p>
Full article ">Figure 2
<p>Computational model and mesh for ignition transience in a solid rocket motor.</p>
Full article ">Figure 3
<p>Pressure change curve with time during ignition and pressure building.</p>
Full article ">Figure 4
<p>Combustion chamber pressure monitoring points.</p>
Full article ">Figure 5
<p>Comparison of pressure–time curves from experiment and numerical simulation for certain type of solid rocket motor.</p>
Full article ">Figure 6
<p>Temporal variation in temperature contours in the combustion chamber before the propellant charge ignited.</p>
Full article ">Figure 7
<p>Temporal variation in temperature contours in the combustion chamber before the cover opened.</p>
Full article ">Figure 8
<p>Temporal variation in temperature contours in the combustion chamber until the burning surface is almost completely ignited.</p>
Full article ">Figure 9
<p>Temporal variation in pressure contours in the combustion chamber.</p>
Full article ">Figure 10
<p>The pressure variation curve of each probing point with ignition time. (<b>a</b>) Pressure versus time curve in the whole time domain. (<b>b</b>) The details of the pressure versus time curve.</p>
Full article ">Figure 11
<p>Combustor profile diagram under the creep condition of solid rocket motor propellant.</p>
Full article ">Figure 12
<p>Temperature distribution of gases in the combustion chamber of solid rocket motors with different storage years at 6ms after ignition.</p>
Full article ">Figure 13
<p>Temperature distribution during the ignition of the propellant column.</p>
Full article ">Figure 14
<p>Temperature distribution when the pressure on the front side of the cover reaches the threshold for opening the cover.</p>
Full article ">Figure 15
<p>The change curve of pressure vs. time. (<b>a</b>) The global pressure change line. (<b>b</b>) The details of pressure variation.</p>
Full article ">Figure 16
<p>The pressure growth comparison curve. (<b>a</b>) The detailed changes in pressure after the cover is opened. (<b>b</b>) The pressure rise rate over time.</p>
Full article ">Figure 17
<p>Temperature distribution and pressure variation curve (point P9) over time for rocket motor after 3 years of storage.</p>
Full article ">
24 pages, 5102 KiB  
Article
The Application of an Abaqus Preprocessor Based on Python Language in a DAHC Negative Poisson Ratio Structure
by Xiaoting Sun, Guibo Yu, Xiujie Zhu, Jinli Che, Yuanyuan Yan and Wei Wang
Crystals 2025, 15(2), 181; https://doi.org/10.3390/cryst15020181 - 13 Feb 2025
Viewed by 362
Abstract
The double-arrow hollow cylinder (DAHC) has excellent mechanical properties under axial compression loads due to its negative Poisson ratio structure. To improve finite element simulation efficiency for parameter optimization analysis, an automatic modeling program for DAHC negative Poisson ratio structures was developed on [...] Read more.
The double-arrow hollow cylinder (DAHC) has excellent mechanical properties under axial compression loads due to its negative Poisson ratio structure. To improve finite element simulation efficiency for parameter optimization analysis, an automatic modeling program for DAHC negative Poisson ratio structures was developed on the ABAQUS platform using Python. The sample was manufactured using selective laser melting additive manufacturing based on the solid model and then simulated using the finite element method. The accuracy of the automatic modeling method was confirmed by comparing load–displacement curves and deformation cloud images with static compression test results. The DAHC automatic modeling program based on Python can reduce workload during parameter optimization analysis and improve finite element analysis efficiency. Full article
(This article belongs to the Section Inorganic Crystalline Materials)
Show Figures

Figure 1

Figure 1
<p>Double-arrow NPR structure and its main parameters.</p>
Full article ">Figure 2
<p>DAHC negative Poisson ratio structure.</p>
Full article ">Figure 3
<p>DAHC negative Poisson ratio structure (Layer 1).</p>
Full article ">Figure 4
<p>Steps of automatic modeling of DAHC negative Poisson ratio structure.</p>
Full article ">Figure 5
<p>3D printer and molded parts.: (<b>a</b>) LCD light curing 3D printer; (<b>b</b>) Main view of the specimens; (<b>c</b>) Top view of the specimen.</p>
Full article ">Figure 6
<p>NPRC compression test: (<b>a</b>) quasi-static compression test; (<b>b</b>) experimental samples on the universal test machine.</p>
Full article ">Figure 7
<p>Grid division results.</p>
Full article ">Figure 8
<p>Deformation is 0 mm.</p>
Full article ">Figure 9
<p>Deformation is 2 mm.</p>
Full article ">Figure 10
<p>Deformation is 4 mm.</p>
Full article ">Figure 11
<p>Deformation is 6 mm.</p>
Full article ">Figure 12
<p>Comparison of results of NPRC compression test and finite element analysis.</p>
Full article ">Figure 13
<p>Stress nephogram at displacement of 4 mm.</p>
Full article ">Figure 14
<p>Stress nephogram at displacement of 6 mm.</p>
Full article ">Figure 15
<p>Principal strain at displacement of 6 mm.</p>
Full article ">
24 pages, 3280 KiB  
Article
Comparative Analysis on Modelling Approaches for the Simulation of Fatigue Disbonding with Cohesive Zone Models
by Johan Birnie, Maria Pia Falaschetti and Enrico Troiani
Aerospace 2025, 12(2), 139; https://doi.org/10.3390/aerospace12020139 - 13 Feb 2025
Viewed by 354
Abstract
Adhesively bonded joints are essential in the aeronautical industry, offering benefits such as weight reduction and enhanced sustainability. However, certifying these joints is challenging due to unreliable methods for assessing their strength and the development of predictive models for fatigue-driven disbonding remains an [...] Read more.
Adhesively bonded joints are essential in the aeronautical industry, offering benefits such as weight reduction and enhanced sustainability. However, certifying these joints is challenging due to unreliable methods for assessing their strength and the development of predictive models for fatigue-driven disbonding remains an ongoing effort. This manuscript presents the implementation and validation of a cohesive zone model for studying high-cycle fatigue disbonding under Mode I and Mixed-Mode loading. The model was integrated into the commercial finite element analysis software Abaqus using user-defined material subroutine (UMAT). Two modelling approaches were investigated: one replacing the adhesive with a cohesive layer, and the other incorporating a cohesive layer at the adhesive’s mid-plane while modelling its entire thickness, using both 2D and 3D techniques. Validation was conducted against experimental data from the literature that examined the influence of adhesive thickness on fatigue behaviour in DCB and CLS tests. The findings of this study confirm that the model accurately predicts fatigue disbonding across all cases examined. Additionally, the analysis reveals that modelling adhesive thickness plays a critical role in the simulation’s outcomes. Variations in adhesive thickness can significantly alter the crack growth behaviour, highlighting the importance of carefully considering this parameter in future assessments and applications. Full article
Show Figures

Figure 1

Figure 1
<p>Visual representation of the bi-linear traction–separation law, non-damaged (<b>a</b>); damaged, illustrating stiffness reduction (<b>b</b>); where <math display="inline"><semantics> <msup> <mi>δ</mi> <mn>0</mn> </msup> </semantics></math> is the onset displacement, <math display="inline"><semantics> <msup> <mi>δ</mi> <mi>f</mi> </msup> </semantics></math> is the propagation displacement, <math display="inline"><semantics> <mi>λ</mi> </semantics></math> is the displacement jump, <span class="html-italic">r</span> is the damage threshold, <math display="inline"><semantics> <msup> <mi>τ</mi> <mn>0</mn> </msup> </semantics></math> is the maximum traction, <span class="html-italic">K</span> is the stiffness of the cohesive zone, <span class="html-italic">d</span> is the damage, and <math display="inline"><semantics> <msub> <mi>G</mi> <mi>c</mi> </msub> </semantics></math> the fracture energy.</p>
Full article ">Figure 2
<p>Fatigue loading diagram, displacement number of cycles, illustrating the displacement’s envelope and ratio (variation) during the fatigue loading phase.</p>
Full article ">Figure 3
<p>Visual representation of the cyclic variation of the strain energy release rate (<math display="inline"><semantics> <mrow> <mi mathvariant="sans-serif">Δ</mi> <mi>G</mi> </mrow> </semantics></math>) of the bi-linear traction separation law, where <math display="inline"><semantics> <msup> <mi>λ</mi> <mrow> <mi>m</mi> <mi>i</mi> <mi>n</mi> </mrow> </msup> </semantics></math> is the minimum and <math display="inline"><semantics> <msup> <mi>λ</mi> <mrow> <mi>m</mi> <mi>a</mi> <mi>x</mi> </mrow> </msup> </semantics></math> the maximum displacement jump during the cyclic loading.</p>
Full article ">Figure 4
<p>Visual representation of the bi-linear traction separation law highlighting in green the equivalent displacement mode for Mixed-Mode loading considering Mode I and Shear.</p>
Full article ">Figure 5
<p>Modelling approaches to the cohesive zone: Cohesive zone replacing the adhesive (<b>a</b>); Cohesive zone located along the mid-plane of the adhesive (<b>b</b>); in green the adherent, in red the adhesive, and in grey the cohesive zone.</p>
Full article ">Figure 6
<p>UMAT implementation workflow for the simulation of high cycle fatigue with cohesive zone models.</p>
Full article ">Figure 7
<p>Visual representation of the DCB during disbonding during fatigue cycling at N = 0 (<b>a</b>); at N = 375,000 (<b>b</b>); damage distribution in cohesive zone at N = 0 (<b>c</b>); and at N = 375,000 (<b>d</b>).</p>
Full article ">Figure 8
<p>Crack growth rate as a function of normalised cyclic variation of the strain energy release rate. Results are divided by specimen name and by the different testing configurations.</p>
Full article ">Figure 9
<p>Crack length as a function of the number of fatigue cycles. Results divided by specimen name and by the different testing configurations.</p>
Full article ">Figure 10
<p>Visual representation of the CLS during disbonding during fatigue cycling at N = 0 (<b>a</b>); at N = 160,000 (<b>b</b>); damage distribution in cohesive zone at N = 0 (<b>c</b>); and at N = 160,000 (<b>d</b>).</p>
Full article ">Figure 11
<p>Crack growth rate as a function of normalised cyclic variation of the strain energy release rate. Results are divided by specimen name and by the different modelling configurations. Solid lines represent the interpolation for each model.</p>
Full article ">Figure 12
<p>Crack length as a function of the number of fatigue cycles. Results divided by the different modelling configurations.</p>
Full article ">
23 pages, 9640 KiB  
Article
Full-Scale Testing and Stability Analysis of Prefabricated Steel Plate-Concrete Composite Walls in Underground Granaries
by Hao Zhang, Ruixin Wang, Lei Chen and Jun Chuai
Buildings 2025, 15(4), 561; https://doi.org/10.3390/buildings15040561 - 12 Feb 2025
Viewed by 370
Abstract
Underground granaries naturally preserve grain quality by maintaining low temperatures and reduced oxygen levels, eliminating the need for artificial cooling and pest control. However, cast-in-place reinforced concrete construction faces challenges such as waterproofing and complex on-site processes, necessitating prefabricated steel plate-concrete composite structures [...] Read more.
Underground granaries naturally preserve grain quality by maintaining low temperatures and reduced oxygen levels, eliminating the need for artificial cooling and pest control. However, cast-in-place reinforced concrete construction faces challenges such as waterproofing and complex on-site processes, necessitating prefabricated steel plate-concrete composite structures with robust joints for enhanced structural integrity and streamlined construction. The study utilizes a full-scale prefabricated steel plate-concrete underground silo, instrumented with strain gauges on circumferential steel bars and internal steel plates to monitor stress variations during six distinct backfilling loading cases. Concurrently, finite element models were developed using ABAQUS 6.14 software for numerical simulations, which were validated against experimental data. Stability analyses, including buckling load assessments and parameter sensitivity studies, were conducted to evaluate the effects of joint quantity and bending stiffness on the structural performance of the composite walls. The results revealed that circumferential joints play a critical role in stress distribution within the composite walls, underscoring the necessity of optimized joint design. The numerical model accurately replicated experimental results, with deviations below 9%, confirming its reliability. Furthermore, an equivalent joint design method was established, demonstrating that a joint bending stiffness ratio above 1.1 ensures that prefabricated composite walls achieve critical buckling loads comparable to cast-in-place walls. These findings provide a robust framework for enhancing the structural performance and reliability of prefabricated underground silos. Full article
(This article belongs to the Section Building Structures)
Show Figures

Figure 1

Figure 1
<p>Prefabricated steel plate-concrete underground grain silo components and construction details.</p>
Full article ">Figure 2
<p>Arrangement of steel bar gauges for measuring circumferential steel in the internal steel plate of the composite silo wall.</p>
Full article ">Figure 3
<p>Arrangement of steel bar gauges for measuring circumferential stress on the internal steel plate of the composite silo wall.</p>
Full article ">Figure 4
<p>Arrangement of earth pressure gauges on the outer joint steel plate of the composite silo wall.</p>
Full article ">Figure 4 Cont.
<p>Arrangement of earth pressure gauges on the outer joint steel plate of the composite silo wall.</p>
Full article ">Figure 5
<p>Schematic diagram of backfilling cases.</p>
Full article ">Figure 6
<p>Earth pressure distribution under different backfilling cases.</p>
Full article ">Figure 7
<p>Circumferential stress value of the circumferential steel bar under different cases.</p>
Full article ">Figure 7 Cont.
<p>Circumferential stress value of the circumferential steel bar under different cases.</p>
Full article ">Figure 8
<p>Circumferential stress value of the internal steel plate under different cases.</p>
Full article ">Figure 9
<p>Circumferential stress value of the internal steel plate at 3.65m.</p>
Full article ">Figure 10
<p>Numerical model of prefabricated steel plate-concrete underground silo.</p>
Full article ">Figure 11
<p>Comparison of simulated and experimental circumferential stress values for the prefabricated composite silo wall.</p>
Full article ">Figure 12
<p>Prefabricated composite silo wall model.</p>
Full article ">Figure 13
<p>First-order buckling mode of the composite silo wall.</p>
Full article ">Figure 14
<p>Curve of buckling load varying with the number of joints.</p>
Full article ">Figure 15
<p>Curve of buckling load varying with the thickness of the internal or outer joint steel plate.</p>
Full article ">Figure 16
<p>Curve of load ratio variation with stiffness ratio for prefabricated composite silo walls.</p>
Full article ">
25 pages, 8099 KiB  
Article
Assessment of Externally Prestressed Beams with FRP Rebars Considering Bond–Slip Effects
by Zhangxiang Li, Bo Chen, Xueliang Wang and Tiejiong Lou
Materials 2025, 18(4), 787; https://doi.org/10.3390/ma18040787 - 11 Feb 2025
Viewed by 462
Abstract
This paper presents detailed numerical modeling of externally prestressed concrete (EPC) beams with fiber-reinforced polymer (FRP) rebars. Particular attention is paid to the bond–slip interactions between FRP rebars and concrete. A refined 3D finite element model (FEM) incorporating a script describing the bond–slip [...] Read more.
This paper presents detailed numerical modeling of externally prestressed concrete (EPC) beams with fiber-reinforced polymer (FRP) rebars. Particular attention is paid to the bond–slip interactions between FRP rebars and concrete. A refined 3D finite element model (FEM) incorporating a script describing the bond–slip of FRP rebars and concrete is developed in ABAQUS. The model effectiveness, rooted in the interface behavior between FRP rebars and concrete, is comprehensively assessed using experimental data. A comprehensive investigation has been conducted using FEM on the mechanical behavior of carbon fiber-reinforced polymer (CFRP) tendon–EPC beams with FRP rebars. Due to the bond–slip effect, FRP rebars in EPC beams exhibit a distinct phenomenon of stress degradation. This suggests that the traditional method based on plane cross-sectional assumptions is no longer suitable for the engineering design of EPC beams with FRP rebars. Moreover, the study assesses several models including typical design codes for their accuracy in predicting the elevation of ultimate stress in external tendons. It is demonstrated that some of the design codes are overly conservative when estimating the tendon stress in EPC beams with FRP rebars. Full article
Show Figures

Figure 1

Figure 1
<p>Stress–strain responses for the materials employed. (<b>a</b>) concrete under compression; (<b>b</b>) concrete under tension; (<b>c</b>) prestressing tendons; (<b>d</b>) bonded rebars.</p>
Full article ">Figure 2
<p>Bond stress–slip graphs of FRP rebars and concrete. (<b>a</b>) ribbed FRP; (<b>b</b>) plain FRP [<a href="#B31-materials-18-00787" class="html-bibr">31</a>,<a href="#B32-materials-18-00787" class="html-bibr">32</a>,<a href="#B36-materials-18-00787" class="html-bibr">36</a>,<a href="#B37-materials-18-00787" class="html-bibr">37</a>].</p>
Full article ">Figure 3
<p>Assignment in mesh attribute and connector of specimens.</p>
Full article ">Figure 4
<p>Comparison with test data for bond–slip model validation. (<b>a</b>) C1; (<b>b</b>) C2 [<a href="#B38-materials-18-00787" class="html-bibr">38</a>].</p>
Full article ">Figure 5
<p>Simple EPC beams. (<b>a</b>) dimensions and rebar details; (<b>b</b>) external tendon configuration; (<b>c</b>) FEM mesh of T-2.</p>
Full article ">Figure 6
<p>Damage nephograms of T-2.</p>
Full article ">Figure 7
<p>Validation against tests. (<b>a</b>) load vs. midspan deflection; (<b>b</b>) load vs. tendon stress [<a href="#B40-materials-18-00787" class="html-bibr">40</a>].</p>
Full article ">Figure 8
<p>Externally prestressed T-beams for numerical assessment.</p>
Full article ">Figure 9
<p>FEM mesh of EPC beams.</p>
Full article ">Figure 10
<p>Nephograms of structure and bond rebars for typical beams.</p>
Full article ">Figure 11
<p>Effect of rebar type. (<b>a</b>) midspan deflection vs. load; (<b>b</b>) load vs. tendon stress increase; (<b>c</b>) tensile rebar strain vs. load.</p>
Full article ">Figure 12
<p>Effect of rebar type. (<b>a</b>) deflection vs. tendon stress increment; (<b>b</b>) moment vs. curvature; (<b>c</b>) moment vs. neutral axis depth.</p>
Full article ">Figure 13
<p>Effect of GFRP rebar elastic modulus. (<b>a</b>) midspan deflection vs. load; (<b>b</b>) load vs. tendon stress increase; (<b>c</b>) tensile rebar strain vs. load.</p>
Full article ">Figure 14
<p>Effect of concrete grade. (<b>a</b>) midspan deflection vs. load; (<b>b</b>) load vs. tendon stress increase; (<b>c</b>) tensile rebar strain vs. load.</p>
Full article ">Figure 15
<p>Effect of concrete grade. (<b>a</b>) deflection vs. tendon stress increment; (<b>b</b>) moment vs. curvature; (<b>c</b>) moment vs. neutral axis depth.</p>
Full article ">Figure 16
<p>Effect of reinforcement ratio on the load-deflection behavior. (<b>a</b>) EPC beams with CFRP rebars; (<b>b</b>) EPC beams with GFRP rebars.</p>
Full article ">Figure 17
<p>Ultimate behavior with varying reinforcement ratio. (<b>a</b>) ultimate deflection; (<b>b</b>) ultimate load; (<b>c</b>) ultimate tendon stress.</p>
Full article ">Figure 18
<p>Ultimate behavior with varying reinforcement ratio. (<b>a</b>) ultimate curvature; (<b>b</b>) ultimate neutral axis depth; (<b>c</b>) ultimate tensile rebar strain.</p>
Full article ">Figure 19
<p>Numerical and code predictions on Δ<span class="html-italic">σ<sub>p</sub></span>. (<b>a</b>) variation in Δ<span class="html-italic">σ<sub>p</sub></span> based on <span class="html-italic">ω</span><sub>0</sub>; (<b>b</b>) variation in Δ<span class="html-italic">σ<sub>p</sub></span> based on <span class="html-italic">c<sub>u</sub></span>/<span class="html-italic">d<sub>p</sub></span> [<a href="#B23-materials-18-00787" class="html-bibr">23</a>,<a href="#B43-materials-18-00787" class="html-bibr">43</a>,<a href="#B44-materials-18-00787" class="html-bibr">44</a>,<a href="#B45-materials-18-00787" class="html-bibr">45</a>,<a href="#B46-materials-18-00787" class="html-bibr">46</a>].</p>
Full article ">Figure 20
<p>Comparison of Δ<span class="html-italic">σ<sub>p</sub></span> by simplified models with numerical predictions [<a href="#B23-materials-18-00787" class="html-bibr">23</a>,<a href="#B43-materials-18-00787" class="html-bibr">43</a>,<a href="#B44-materials-18-00787" class="html-bibr">44</a>,<a href="#B45-materials-18-00787" class="html-bibr">45</a>,<a href="#B46-materials-18-00787" class="html-bibr">46</a>].</p>
Full article ">
21 pages, 10288 KiB  
Article
Finite Element Modeling of Dynamic Response of RPC Columns and Frames Under Coupled Fire and Explosion
by Qin Rong, Chaochao Peng, Xiaomeng Hou, Yuan Chang and Tiancong Fan
Appl. Sci. 2025, 15(3), 1668; https://doi.org/10.3390/app15031668 - 6 Feb 2025
Viewed by 563
Abstract
Reactive powder concrete (RPC) is widely used in ultra-high-rise buildings, hydropower stations, bridges, and other important infrastructures. To study the dynamic response and damage characteristics of RPC columns and frames considering coupled fire and explosions, an analytical model of RPC columns and frames [...] Read more.
Reactive powder concrete (RPC) is widely used in ultra-high-rise buildings, hydropower stations, bridges, and other important infrastructures. To study the dynamic response and damage characteristics of RPC columns and frames considering coupled fire and explosions, an analytical model of RPC columns and frames with coupled fire and explosions was established by using ABAQUS (2021) finite element software. The dynamic response and damage degree of RPC columns under coupled fire and explosions were investigated to reveal the influence laws of parameters such as cross-section size, axial compression ratio, reinforcement rate, and fire duration on the dynamic response of RPC columns at high temperatures. The dynamic response of the frame structure was analyzed when the explosion load was applied to the bottom corner columns, side columns, and top beams, respectively. The results show that the fire severely weakened the blast resistance of RPC columns; the maximum mid-span deformation and residual deformation of RPC columns decreased with the increase in cross-section size and longitudinal bar reinforcement ratio and increased with the increase in fire duration and axial compression ratio. When the explosion load was applied to the corner columns of the bottom floor of the frame, the bottom corner columns were almost completely destroyed, and there was a significant risk of the structure collapsing. Based on the results of the data analysis, a method to enhance the explosion resistance of RC frame structures using RPC materials at high temperatures is proposed. Full article
(This article belongs to the Special Issue Emerging Technologies of Sustainable Building Materials)
Show Figures

Figure 1

Figure 1
<p>Comparison between experimental and simulated values of temperature field.</p>
Full article ">Figure 2
<p>Column geometry and reinforcement layout.</p>
Full article ">Figure 3
<p>Time history curve of RPC column mid-span deformation.</p>
Full article ">Figure 4
<p>Comparison between simulated stress–strain curve results and measured results.</p>
Full article ">Figure 5
<p>RPC column numerical model.</p>
Full article ">Figure 6
<p>Horizontal displacement–time course curves in RPC column spans with different fire times.</p>
Full article ">Figure 7
<p>Plot of horizontal displacement in RPC column span with fire time.</p>
Full article ">Figure 8
<p>Time–distance curves of horizontal displacements in RPC column spans of different cross-section sizes.</p>
Full article ">Figure 9
<p>Plot of horizontal displacement in RPC column span as a function of section size.</p>
Full article ">Figure 10
<p>Time course of horizontal displacement in span of RPC columns with different axial pressure ratios.</p>
Full article ">Figure 11
<p>Plot of horizontal displacement with axial pressure ratio in RPC column span.</p>
Full article ">Figure 12
<p>Horizontal displacement–time course curves of columns in RPC columns span under columns with different reinforcement ratios.</p>
Full article ">Figure 13
<p>Plot of horizontal displacement with reinforcement ratio in RPC column span.</p>
Full article ">Figure 14
<p>Concrete frame numerical model.</p>
Full article ">Figure 15
<p>Explosion load located on a corner pillar on the first floor.</p>
Full article ">Figure 16
<p>Explosion load located on the first layer edge pillar.</p>
Full article ">Figure 17
<p>Explosion load located at the top of a beam on the first floor.</p>
Full article ">Figure 18
<p>Diagram of frame damage when explosion load located on a corner pillar on the first floor.</p>
Full article ">Figure 19
<p>Diagram of frame damage when explosion load located on the first layer edge pillar.</p>
Full article ">Figure 20
<p>Diagram of frame damage when explosion load located at the top of a beam on the first floor.</p>
Full article ">
18 pages, 10071 KiB  
Article
Crack Propagation in Axial-Flow Fan Blades Under Complex Loading Conditions: A FRANC3D and ABAQUS Co-Simulation Approach
by Mariem Ben Hassen, Slim Ben-Elechi and Hatem Mrad
Appl. Sci. 2025, 15(3), 1597; https://doi.org/10.3390/app15031597 - 5 Feb 2025
Viewed by 430
Abstract
Since fan blades are exposed to fatigue, and in some cases harsh loading conditions, they may exhibit fracture failures due to crack propagation, resulting in significant losses. Previous studies of crack propagation in blades are mainly confined to either simplified blade geometry or [...] Read more.
Since fan blades are exposed to fatigue, and in some cases harsh loading conditions, they may exhibit fracture failures due to crack propagation, resulting in significant losses. Previous studies of crack propagation in blades are mainly confined to either simplified blade geometry or loads, resulting in a significant discrepancy between the simulated crack propagation and the real blade propagation behavior, while it is lacking for challenging shapes and loads. A co-simulation approach of FRANC3D and ABAQUS was developed to study the crack propagation of an axial-flow fan blade subjected to centrifugal, aerodynamic, and combined loads. The projected approach is validated with results obtained from analytical calculations and experiments. Meanwhile, making use of benchmarks, the Stress Intensity Factor (SIF) and the prediction of mixed-mode crack growth path are validated. Considering various loads, the crack propagation path response for the fan blade is computed for different growth steps. The results pinpoint that the crack propagation length of the crack tip center is maximum under centrifugal loading. However, the aerodynamic load led to a maximum propagation length of the crack tip endpoints. In addition, the combined force of centrifugal and aerodynamic loads limits the crack from growing. Full article
Show Figures

Figure 1

Figure 1
<p>Finite element model illustration of the specimen: (<b>a</b>) without crack, (<b>b</b>) with center double slant crack.</p>
Full article ">Figure 2
<p>Comparison of the theoretical and numerical results of stress intensity factors: KI and KII with different initial crack lengths: (<b>a</b>) for φ = 30°, and (<b>b</b>) for φ = 45°.</p>
Full article ">Figure 3
<p>The polymethyl methacrylate (PMMA) beam specimens: (<b>a</b>) the finite element model without holes, (<b>b</b>) the crack growth path for specimen A, (<b>c</b>) the crack growth path for specimen B, (<b>d</b>) the crack growth path for specimen C.</p>
Full article ">Figure 4
<p>The PMMA specimen: (<b>a</b>) the finite element model with holes, (<b>b</b>) the crack growth path for specimen D, (<b>c</b>) the crack growth path for specimen E, (<b>d</b>) the crack growth path for specimen F.</p>
Full article ">Figure 5
<p>Comparison of the crack propagation paths obtained using the proposed method with the experimental [<a href="#B24-applsci-15-01597" class="html-bibr">24</a>] and Abaqus self-method for the first three PMMA beam specimens (A, B, C).</p>
Full article ">Figure 6
<p>Comparison of the crack propagation paths obtained using the proposed method with the experimental [<a href="#B24-applsci-15-01597" class="html-bibr">24</a>] and the Abaqus self-method for the last three PMMA beam specimens (D, E, F).</p>
Full article ">Figure 7
<p>Crack propagation flow chart using FRANC3D code.</p>
Full article ">Figure 8
<p>Boundary conditions acting on the blade: (<b>a</b>) centrifugal load, (<b>b</b>) aerodynamic load, (<b>c</b>) combined centrifugal and aerodynamic loads.</p>
Full article ">Figure 9
<p>Equivalent (von Mises) stress distribution along the blade: (<b>a</b>) centrifugal load, (<b>b</b>) aerodynamic load, (<b>c</b>) combined centrifugal and aerodynamic loads.</p>
Full article ">Figure 10
<p>Crack propagation path under: (<b>a</b>) centrifugal load, (<b>b</b>) aerodynamic load, (<b>c</b>) combined centrifugal and aerodynamic loads.</p>
Full article ">Figure 11
<p>Stress intensity factor KI for: (<b>a</b>) centrifugal load, (<b>b</b>) aerodynamic load, (<b>c</b>) combined centrifugal and aerodynamic loads.</p>
Full article ">Figure 12
<p>The crack propagation lengths of both ends of the crack tip for the three equal subdivisions of the step number.</p>
Full article ">Figure 13
<p>The crack propagation lengths of the center of the crack tip for the three equal subdivisions of the step number.</p>
Full article ">Figure 14
<p>Comparison of the total crack length for the different loads.</p>
Full article ">
Back to TopTop