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Search Results (593)

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19 pages, 1228 KiB  
Article
Stair-Climbing Wheeled Robot Based on Rotating Locomotion of Curved-Spoke Legs
by Dongwoo Seo and Jaeyoung Kang
Biomimetics 2024, 9(10), 633; https://doi.org/10.3390/biomimetics9100633 - 17 Oct 2024
Abstract
This study proposes a new wheel-leg mechanism concept and formulations for the kinematics and dynamics of a stair-climbing robot utilizing the rotating leg locomotion of curved spokes and rolling tires. The system consists of four motor-driven tires and four curved-spoke legs. The curved-spoke [...] Read more.
This study proposes a new wheel-leg mechanism concept and formulations for the kinematics and dynamics of a stair-climbing robot utilizing the rotating leg locomotion of curved spokes and rolling tires. The system consists of four motor-driven tires and four curved-spoke legs. The curved-spoke leg is semicircle-like and is used to climb stairs. Once the spoke leg rolls on the surface, it lifts and pulls the mating wheel toward the surface, owing to the kinematic constraint between the spoke and the wheel. Single-wheel climbing is a necessary condition for the stair climbing of whole robots equipped with front and rear axles. This study proposes the design requirements of a spoke leg for the success of single-wheel climbing in terms of kinematic inequality equations according to the scenario of single-wheel climbing. For a design configuration that enables single-wheel climbing, the required minimum friction coefficient for the static analysis of the stair-climbing wheeled robots is demonstrated. Thereafter, the stair-climbing ability is validated through the dynamic equations that enable the frictional slip of the tires, as well as the curved-spoke legs. Lastly, the results revealed that the rotating locomotion of the well-designed curved-spoke legs effectively enables the stair climbing of the whole robot. Full article
(This article belongs to the Special Issue Design and Control of a Bio-Inspired Robot: 3rd Edition)
23 pages, 12281 KiB  
Article
Research on the Hydrodynamic Characteristics of a Rectangular Otter Board in Different Work Postures Based on a Dynamic Model
by Wenhua Chu, Minghao Zhai, Senqi Cui, Yu Cao, Xinyang Zhang and Qiaoli Zhou
J. Mar. Sci. Eng. 2024, 12(10), 1856; https://doi.org/10.3390/jmse12101856 - 17 Oct 2024
Abstract
This paper investigates the hydrodynamic characteristics of a rectangular otter board in different working postures by using a dynamic model. Dynamic models are mainly based on dynamic mesh techniques. The results of the dynamic model are, compared to the model test, carried out [...] Read more.
This paper investigates the hydrodynamic characteristics of a rectangular otter board in different working postures by using a dynamic model. Dynamic models are mainly based on dynamic mesh techniques. The results of the dynamic model are, compared to the model test, carried out in a flume tank. Furthermore, different rotation speeds of dynamic model were analyzed. The research results are as follows: compared to flume tank results, the maximum error of the dynamic model is 23.77%. Moreover, the influence of rotation speed on the hydrodynamic board is not obvious, and 2 deg./s was chosen as the rotation speed. When the board is tilted slightly (including four working postures), its lift-to-drag ratio first increases slightly and then gradually decreases. Compared with the other three working postures, the pressure center coefficient of the board does not change significantly when it is tilted inward. When studying different working angles (including AOA and tilt angle) of the otter board, the numerical dynamic model significantly reduces repetitive setup work, making simulations more efficient. Its ability to provide continuous curves and a large volume of results offers researchers a more detailed and comprehensive understanding of the board’s hydrodynamics. Additionally, the dynamic model supports innovative fishery equipment development by allowing more accurate and continuous numerical simulations. Full article
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Figure 1
<p>Diagram of rectangular otter board model (<b>a</b>) Front view (<b>b</b>) Rear view (<b>c</b>) Three-dimensional view.</p>
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<p>Different working postures of the otter board.</p>
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<p>Numerical simulation calculation area.</p>
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<p>Mesh division (AOA = 0°).</p>
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<p>Rectangular otter board model test design schematic diagram.</p>
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<p>The process of rectangular otter board model test.</p>
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<p>Moment and force of board.</p>
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<p>Comparison of (<b>a</b>) lift coefficient and (<b>b</b>) moment coefficient between numerical dynamic model and model test. (<b>c</b>) Comparison of drag or lift coefficient of numerical dynamic model, numerical static model and model test.</p>
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<p>Comparison of (<b>a</b>) lift coefficient and (<b>b</b>) moment coefficient between numerical dynamic model and model test. (<b>c</b>) Comparison of drag or lift coefficient of numerical dynamic model, numerical static model and model test.</p>
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<p>The lift coefficient of the board at different speeds.</p>
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<p>The continuous change in lift coefficient with tilt angle (forward and backward).</p>
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<p>The continuous change in drag coefficient with tilt angle (forward and backward).</p>
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<p>The continuous change in lift-to-drag ratio with tilt angles (forward and backward).</p>
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<p>The continuous change of C<sub>pb</sub> with tilt attitude (forward and backward).</p>
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<p>The continuous change of C<sub>pc</sub> with tilt attitude (forward and backward).</p>
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<p>The continuous change in M<sub>y</sub> with tilt attitude (forward and backward).</p>
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<p>Pressure distribution on the surface of the board (forward).</p>
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<p>Pressure distribution on the surface of the board (backward).</p>
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<p>Distribution of flow field around the board in different tilt forward angles.</p>
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<p>Distribution of flow field around the board in different tilt backward angles.</p>
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<p>The continuous change in lift coefficient with tilt angles (inward and outward).</p>
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<p>The continuous change in drag coefficient with tilt angles (inward and outward).</p>
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<p>The continuous change in the lift-to-drag ratio with tilt angles (inward and outward).</p>
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<p>The continuous change of C<sub>pb</sub> with tilt angle (inward and outward).</p>
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<p>The continuous change of C<sub>pc</sub> with tilt angle (inward and outward).</p>
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<p>Pressure distribution on the surface of the board (inward).</p>
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<p>Pressure distribution on the surface of the board (outward).</p>
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<p>Distribution of flow field around the board in different tilt inward angle.</p>
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<p>Distribution of flow field around the board in different tilt outward angle.</p>
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19 pages, 9442 KiB  
Article
Optimal Selection of Active Jet Parameters for a Ducted Tail Wing Aimed at Improving Aerodynamic Performance
by Huayu Jia, Huilong Zheng, Hong Zhou and Shunbo Huo
Aerospace 2024, 11(10), 851; https://doi.org/10.3390/aerospace11100851 - 15 Oct 2024
Viewed by 241
Abstract
The foldable tail of the box-type launch vehicle poses a risk of mechanical jamming during the launch process, which is not conducive to the smooth completion of the flight mission. The integrated nonfolding ducted tail proposed in this article can solve the problem [...] Read more.
The foldable tail of the box-type launch vehicle poses a risk of mechanical jamming during the launch process, which is not conducive to the smooth completion of the flight mission. The integrated nonfolding ducted tail proposed in this article can solve the problem of storing the tail in the launch box. Moreover, traditional mechanical control surfaces have been eliminated, and active jet control has been adopted to control the pitch direction of the flight attitude, which can improve the structural reliability of the tail wing. By studying the effects of parameters such as momentum coefficient, jet hole position, jet hole height, and jet angle on improving the aerodynamic performance of ducted tail wing, relatively good jet parameters are selected. Research has found that compared with jet hole height and jet angle, momentum coefficient and jet hole position are more effective in improving the aerodynamic performance of ducted tail wings. Under a trailing edge jet, a relatively good jet condition occurs when the jet hole height is equal to0.25% of the aerodynamic chord length, and the jet angle is equal to 0°. At this time, with the increase of the jet momentum coefficient, the effect of increasing the lift of the ducted tail wing is the best. Finally, a comparative analysis is conducted on the lift and drag characteristics between the ducted tail wing and traditional tail wing, and it is found that the ducted tail wing can generate lift at a 0° attack angle and will not stall in the high attack angle range of 12°~22°, with broad application prospects. Full article
(This article belongs to the Special Issue Aerodynamic Numerical Optimization in UAV Design)
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Figure 1
<p>The advantages and parametric modeling of active jet ducted tail fins. (<b>a</b>) Compared with the traditional tail, the ducted tail in the launch box does not need to be folded; (<b>b</b>) during flight, the aircraft controls its flight attitude through an active jet ducted tail fin; (<b>c</b>) modeling jet parameters for ducted tail fins.</p>
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<p>Details of the grid. (<b>a</b>) Grid of ordinary tail wing; (<b>b</b>) grid of jet ducted tail wing.</p>
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<p>Verification of grid independence for the ordinary NACA0012 tail wing.</p>
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<p>Verification of calculation accuracy for the ordinary NACA0012 tail wing.</p>
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<p>The improvement effect of momentum coefficient on the aerodynamic performance under trailing edge jet conditions. (<b>a</b>) Lift coefficient; (<b>b</b>) growth value of lift coefficient; (<b>c</b>) drag coefficient; (<b>d</b>) growth value of drag coefficient.</p>
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<p>The improvement effect of momentum coefficient on the aerodynamic performance under 20% leading edge jet conditions. (<b>a</b>) Lift coefficient; (<b>b</b>) growth value of lift coefficient; (<b>c</b>) drag coefficient; (<b>d</b>) growth value of drag coefficient.</p>
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<p>The improvement effect of momentum coefficient on the aerodynamic performance under 20% leading edge jet conditions. (<b>a</b>) Lift coefficient; (<b>b</b>) growth value of lift coefficient; (<b>c</b>) drag coefficient; (<b>d</b>) growth value of drag coefficient.</p>
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<p>The improvement effect of jet hole height on aerodynamic performance in the state of momentum coefficient 0.02. (<b>a</b>) Lift coefficient; (<b>b</b>) growth value of lift coefficient; (<b>c</b>) drag coefficient; (<b>d</b>) growth value of drag coefficient.</p>
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<p>The improvement effect of jet hole height on aerodynamic performance in the state of momentum coefficient 0.04. (<b>a</b>) Lift coefficient; (<b>b</b>) growth value of lift coefficient; (<b>c</b>) drag coefficient; (<b>d</b>) growth value of drag coefficient.</p>
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<p>The improvement effect of jet hole height on aerodynamic performance in the state of momentum coefficient 0.04. (<b>a</b>) Lift coefficient; (<b>b</b>) growth value of lift coefficient; (<b>c</b>) drag coefficient; (<b>d</b>) growth value of drag coefficient.</p>
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<p>The improvement effect of jet hole position on aerodynamic performance at a jet angle of 0 ° and momentum coefficient of 0.04. (<b>a</b>) Lift coefficient; (<b>b</b>) growth value of lift coefficient; (<b>c</b>) drag coefficient; (<b>d</b>) growth value of drag coefficient.</p>
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<p>The improvement effect of jet hole position on aerodynamic performance at a jet angle of 10 ° and momentum coefficient of 0.04. (<b>a</b>) Lift coefficient; (<b>b</b>) growth value of lift coefficient; (<b>c</b>) drag coefficient; (<b>d</b>) growth value of drag coefficient.</p>
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<p>The improvement effect of jet angle on aerodynamic performance under the condition of leading edge 20% and momentum coefficient 0.04. (<b>a</b>) Lift coefficient; (<b>b</b>) growth value of lift coefficient; (<b>c</b>) drag coefficient; (<b>d</b>) growth value of drag coefficient.</p>
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<p>The improvement effect of jet angle on aerodynamic performance under the condition of leading edge 40% and momentum coefficient 0.04. (<b>a</b>) Lift coefficient; (<b>b</b>) growth value of lift coefficient; (<b>c</b>) drag coefficient; (<b>d</b>) growth value of drag coefficient.</p>
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<p>Comparison of lift and drag characteristics between the ducted tail and conventional tail. (<b>a</b>) Comparison of lift coefficients; (<b>b</b>) comparison of drag coefficients.</p>
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<p>Cloud maps of ducted tail wing. (<b>a</b>) <math display="inline"><semantics> <mrow> <msub> <mi>C</mi> <mi>μ</mi> </msub> <mo>=</mo> <mn>0.01</mn> <mo>,</mo> <mi>α</mi> <mo>=</mo> <mn>0</mn> <mo>°</mo> </mrow> </semantics></math> (<b>b</b>) <math display="inline"><semantics> <mrow> <msub> <mi>C</mi> <mi>μ</mi> </msub> <mo>=</mo> <mn>0.04</mn> <mo>,</mo> <mi>α</mi> <mo>=</mo> <mn>0</mn> <mo>°</mo> </mrow> </semantics></math>; (<b>c</b>) <math display="inline"><semantics> <mrow> <msub> <mi>C</mi> <mi>μ</mi> </msub> <mo>=</mo> <mn>0.01</mn> <mo>,</mo> <mi>α</mi> <mo>=</mo> <mn>20</mn> <mo>°</mo> </mrow> </semantics></math>; (<b>d</b>) <math display="inline"><semantics> <mrow> <msub> <mi>C</mi> <mi>μ</mi> </msub> <mo>=</mo> <mn>0.04</mn> <mo>,</mo> <mi>α</mi> <mo>=</mo> <mn>20</mn> <mo>°</mo> </mrow> </semantics></math>.</p>
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<p>Pressure cloud maps for ducted tail wing. (<b>a</b>) <math display="inline"><semantics> <mrow> <msub> <mi>C</mi> <mi>μ</mi> </msub> <mo>=</mo> <mn>0.01</mn> <mo>,</mo> <mi>α</mi> <mo>=</mo> <mn>0</mn> <mo>°</mo> </mrow> </semantics></math> (<b>b</b>) <math display="inline"><semantics> <mrow> <msub> <mi>C</mi> <mi>μ</mi> </msub> <mo>=</mo> <mn>0.04</mn> <mo>,</mo> <mi>α</mi> <mo>=</mo> <mn>0</mn> <mo>°</mo> </mrow> </semantics></math>; (<b>c</b>) <math display="inline"><semantics> <mrow> <msub> <mi>C</mi> <mi>μ</mi> </msub> <mo>=</mo> <mn>0.01</mn> <mo>,</mo> <mi>α</mi> <mo>=</mo> <mn>20</mn> <mo>°</mo> </mrow> </semantics></math>; (<b>d</b>) <math display="inline"><semantics> <mrow> <msub> <mi>C</mi> <mi>μ</mi> </msub> <mo>=</mo> <mn>0.04</mn> <mo>,</mo> <mi>α</mi> <mo>=</mo> <mn>20</mn> <mo>°</mo> </mrow> </semantics></math>.</p>
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<p>Comparison of pressure coefficients of ducted tail wing. (<b>a</b>) Comparison of pressure coefficients at <math display="inline"><semantics> <mrow> <mi>α</mi> <mo>=</mo> <mn>0</mn> <mo>°</mo> </mrow> </semantics></math> (<b>b</b>) comparison of pressure coefficients at <math display="inline"><semantics> <mrow> <mi>α</mi> <mo>=</mo> <mn>20</mn> <mo>°</mo> </mrow> </semantics></math>.</p>
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19 pages, 6164 KiB  
Article
Calibration of Discrete Element Method Parameters for a High-Fidelity Lunar Regolith Simulant Considering the Effects of Realistic Particle Shape
by Ningxi Zhou, Jian Chen, Ning Tian, Kaiwei Tian, Juehao Huang and Peng Wu
Materials 2024, 17(19), 4789; https://doi.org/10.3390/ma17194789 - 29 Sep 2024
Viewed by 355
Abstract
The Discrete Element Method (DEM) is an important tool for investigating the geotechnical properties of lunar regolith. The accuracy of DEM simulations largely depends on precise particle modeling and the appropriate selection of mesoscopic parameters. To enhance the reliability and accuracy of the [...] Read more.
The Discrete Element Method (DEM) is an important tool for investigating the geotechnical properties of lunar regolith. The accuracy of DEM simulations largely depends on precise particle modeling and the appropriate selection of mesoscopic parameters. To enhance the reliability and accuracy of the DEM in lunar regolith studies, this paper utilized the high-fidelity IRSM-1 lunar regolith simulant to construct a DEM model with realistic particle shapes and conducted an angle of repose (AoR) simulation test. The optimal DEM parameters were calibrated using a combination of the Plackett–Burman test, steepest ascent test, and Box–Behnken design. The results indicate that the sliding friction coefficient, rolling friction coefficient, and surface energy significantly influence the simulation AoR. By optimizing against the measured AoR using a second-order regression model, the optimal parameter values were determined to be 0.633, 0.401, and 0.2, respectively. Under these optimal parameters, the error between the simulation and experimental AoR was 2.1%. Finally, the calibrated mesoscopic parameters were validated through a lifting cylinder test, showing an error of 6.3% between the simulation and experimental results. The high similarity in the shape of the AoR further confirms the accuracy and reliability of the parameter calibration method. This study provides a valuable reference for future DEM-based research on the mechanical and engineering properties of lunar regolith. Full article
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<p>Appearance (<b>a</b>) and micro morphology (<b>b</b>) of the IRSM-1 lunar regolith.</p>
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<p>Comparison of the chemical composition of IRSM-1 with Apollo lunar regolith samples.</p>
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<p>Mineralogical analysis of IRSM-1 lunar regolith.</p>
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<p>Comparison of the grain size distribution curves between IRSM-1 and other samples.</p>
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<p>Contour coordinate extraction process of IRSM-1 particle: (<b>a</b>) Import the image. (<b>b</b>) Grayscale processing. (<b>c</b>) Threshold segmentation. (<b>d</b>) Contour extraction.</p>
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<p>Definition and calculation of particle shape parameters.</p>
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<p>Angle of repose experiment of IRSM-1 lunar regolith simulant.</p>
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<p>Comparison of clumps generated under different parameter combinations.</p>
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<p>Grain size distribution curves used in DEM model.</p>
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<p>DEM simulation test of repose angle.</p>
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<p>Particle shape characteristics of IRSM-1. (<b>a</b>) Cumulative average value, (<b>b</b>) cumulative distribution.</p>
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<p>Influence of interaction terms on angle of repose: (<b>a</b>) Interaction between <span class="html-italic">μ</span><sub>1</sub> and <span class="html-italic">μ</span><sub>r1</sub>. (<b>b</b>) Interaction between <span class="html-italic">μ</span><sub>1</sub> and <span class="html-italic">γ</span>. (<b>c</b>) Interaction between <span class="html-italic">μ</span><sub>r1</sub> and <span class="html-italic">γ</span>.</p>
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<p>Comparison of the AoR results between simulation and physical tests: (<b>a</b>) The simulation test, (<b>b</b>) the physical test.</p>
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<p>Simulation process of the lifting cylinder test: (<b>a</b>) Particle generation, (<b>b</b>) settling, (<b>c</b>) cylinder lifting, (<b>d</b>) repose angle formation.</p>
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<p>Results of the lifting cylinder test: (<b>a</b>) The simulation test, (<b>b</b>) the physical test.</p>
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19 pages, 6931 KiB  
Article
The Effect of Reynolds Numbers on Flow-Induced Vibrations: A Numerical Study of a Cylinder on Elastic Supports
by Chunhui Ma, Fenglai Huang, Bin Li, Xujian Li and Yu Liu
Water 2024, 16(19), 2765; https://doi.org/10.3390/w16192765 - 28 Sep 2024
Viewed by 554
Abstract
In the field of fluid dynamics, the Reynolds number is a key parameter that influences the flow characteristics around bluff bodies. While its impact on flow around stationary cylinders has been extensively studied, systematic research into flow-induced vibrations (FIVs) under these conditions remains [...] Read more.
In the field of fluid dynamics, the Reynolds number is a key parameter that influences the flow characteristics around bluff bodies. While its impact on flow around stationary cylinders has been extensively studied, systematic research into flow-induced vibrations (FIVs) under these conditions remains limited. This study utilizes numerical simulations to explore the FIV characteristics of smooth cylinders and passive turbulence control (PTC) cylinders supported elastically within a Reynolds number range from 0.8 × 104 to 1.1 × 105. By comparing the vibration responses, lift coefficients, and wake structures of these cylinders across various Reynolds numbers, this paper aims to elucidate how Reynolds numbers affect the flow and vibration characteristics of these structures. The research employs images of instantaneous lift changes and vortex shedding across multiple sections to visually demonstrate the dynamic changes in flow states. The findings are expected to provide theoretical support for optimizing structural design and vibration control strategies in high-Reynolds-number environments, emphasizing the importance of considering Reynolds numbers in structural safety and design optimization. Full article
(This article belongs to the Special Issue Design and Optimization of Fluid Machinery, 2nd Edition)
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Figure 1
<p>Computational domain and geometric model.</p>
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<p>Computational mesh: (<b>a</b>) global view; (<b>b</b>) smooth cylinder close-up view; (<b>c</b>) P5-60 close-up view.</p>
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<p>Vibration responses of smooth cylinder [<a href="#B32-water-16-02765" class="html-bibr">32</a>].</p>
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<p>Lissajous figures with different reduced velocities: (<b>a</b>) experiment results [<a href="#B32-water-16-02765" class="html-bibr">32</a>]; (<b>b</b>) present study.</p>
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<p>Vibration responses of smooth cylinder versus reduced velocity.</p>
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<p>Peak amplitude ratio versus ratio of <math display="inline"><semantics> <mrow> <msubsup> <mrow> <mi>A</mi> </mrow> <mrow> <mi>m</mi> <mi>a</mi> <mi>x</mi> </mrow> <mrow> <mi mathvariant="normal">*</mi> </mrow> </msubsup> </mrow> </semantics></math> to Re.</p>
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<p>Lift coefficients and phase differences.</p>
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<p>End lift correlation versus reduced velocities of smooth cylinders.</p>
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<p>Instantaneous lift contours of smooth cylinder: (<b>a</b>) <math display="inline"><semantics> <mrow> <msup> <mrow> <mi>U</mi> </mrow> <mrow> <mi mathvariant="normal">*</mi> </mrow> </msup> <mo>=</mo> <mn>3.6</mn> </mrow> </semantics></math>, K30-smooth; (<b>b</b>) <math display="inline"><semantics> <mrow> <msup> <mrow> <mi>U</mi> </mrow> <mrow> <mi mathvariant="normal">*</mi> </mrow> </msup> <mo>=</mo> <mn>6.6</mn> </mrow> </semantics></math>, K60-smooth; (<b>c</b>) <math display="inline"><semantics> <mrow> <msup> <mrow> <mi>U</mi> </mrow> <mrow> <mi mathvariant="normal">*</mi> </mrow> </msup> <mo>=</mo> <mn>8.7</mn> </mrow> </semantics></math>, K120-smooth; (<b>d</b>) <math display="inline"><semantics> <mrow> <msup> <mrow> <mi>U</mi> </mrow> <mrow> <mi mathvariant="normal">*</mi> </mrow> </msup> <mo>=</mo> <mn>10.7</mn> </mrow> </semantics></math>, K240-smooth.</p>
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<p>Wake structures at different cross-sections of smooth cylinder: (<b>a</b>) <math display="inline"><semantics> <mrow> <msup> <mrow> <mi>U</mi> </mrow> <mrow> <mi mathvariant="normal">*</mi> </mrow> </msup> <mo>=</mo> <mn>3.6</mn> </mrow> </semantics></math>, K30-smooth; (<b>b</b>) <math display="inline"><semantics> <mrow> <msup> <mrow> <mi>U</mi> </mrow> <mrow> <mi mathvariant="normal">*</mi> </mrow> </msup> <mo>=</mo> <mn>6.6</mn> </mrow> </semantics></math>, K60-smooth; (<b>c</b>) <math display="inline"><semantics> <mrow> <msup> <mrow> <mi>U</mi> </mrow> <mrow> <mi mathvariant="normal">*</mi> </mrow> </msup> <mo>=</mo> <mn>8.7</mn> </mrow> </semantics></math>, K120-smooth; (<b>d</b>) <math display="inline"><semantics> <mrow> <msup> <mrow> <mi>U</mi> </mrow> <mrow> <mi mathvariant="normal">*</mi> </mrow> </msup> <mo>=</mo> <mn>10.7</mn> </mrow> </semantics></math>, K240-smooth.</p>
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<p>Vibration responses of PTC cylinder versus reduced velocity.</p>
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<p>Lift coefficients and phase differences of PTC cylinders.</p>
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<p>End lift correlation versus reduced velocities of PTC cylinders.</p>
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<p>Instantaneous lift contours of PTC cylinder: (<b>a</b>) <math display="inline"><semantics> <mrow> <msup> <mrow> <mi>U</mi> </mrow> <mrow> <mi mathvariant="normal">*</mi> </mrow> </msup> <mo>=</mo> <mn>3.6</mn> </mrow> </semantics></math>, K30-PTC; (<b>b</b>) <math display="inline"><semantics> <mrow> <msup> <mrow> <mi>U</mi> </mrow> <mrow> <mi mathvariant="normal">*</mi> </mrow> </msup> <mo>=</mo> <mn>6.6</mn> </mrow> </semantics></math>, K60-PTC; (<b>c</b>) <math display="inline"><semantics> <mrow> <msup> <mrow> <mi>U</mi> </mrow> <mrow> <mi mathvariant="normal">*</mi> </mrow> </msup> <mo>=</mo> <mn>8.7</mn> </mrow> </semantics></math>, K120-PTC; (<b>d</b>) <math display="inline"><semantics> <mrow> <msup> <mrow> <mi>U</mi> </mrow> <mrow> <mi mathvariant="normal">*</mi> </mrow> </msup> <mo>=</mo> <mn>10.7</mn> </mrow> </semantics></math>, K240-PTC.</p>
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<p>Wake structures at different cross-sections of PTC cylinder: (<b>a</b>) <math display="inline"><semantics> <mrow> <msup> <mrow> <mi>U</mi> </mrow> <mrow> <mi mathvariant="normal">*</mi> </mrow> </msup> <mo>=</mo> <mn>3.6</mn> </mrow> </semantics></math>, K30-PTC; (<b>b</b>) <math display="inline"><semantics> <mrow> <msup> <mrow> <mi>U</mi> </mrow> <mrow> <mi mathvariant="normal">*</mi> </mrow> </msup> <mo>=</mo> <mn>6.6</mn> </mrow> </semantics></math>, K60-PTC; (<b>c</b>) <math display="inline"><semantics> <mrow> <msup> <mrow> <mi>U</mi> </mrow> <mrow> <mi mathvariant="normal">*</mi> </mrow> </msup> <mo>=</mo> <mn>8.7</mn> </mrow> </semantics></math>, K120-PTC; (<b>d</b>) <math display="inline"><semantics> <mrow> <msup> <mrow> <mi>U</mi> </mrow> <mrow> <mi mathvariant="normal">*</mi> </mrow> </msup> <mo>=</mo> <mn>10.7</mn> </mrow> </semantics></math>, K240-PTC.</p>
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22 pages, 11135 KiB  
Article
Assessment of Breakwater as a Protection System against Aerodynamic Loads Acting on the Floating PV System
by Balram Panjwani
Energies 2024, 17(19), 4873; https://doi.org/10.3390/en17194873 - 28 Sep 2024
Viewed by 391
Abstract
Offshore floating photovoltaic (FPV) systems are subjected to significant aerodynamic forces, especially during extreme wind conditions. Accurate estimation of these forces is crucial for the proper design of mooring lines and connection systems. In this study, detailed CFD simulations were performed for various [...] Read more.
Offshore floating photovoltaic (FPV) systems are subjected to significant aerodynamic forces, especially during extreme wind conditions. Accurate estimation of these forces is crucial for the proper design of mooring lines and connection systems. In this study, detailed CFD simulations were performed for various PV panel configurations, and using these CFD simulation correlations were developed to estimate lift and drag forces as a function of the number of panels. These correlations provide valuable tools for designing large-scale FPV systems with multiple PV modules. Additionally, this study investigates the potential of using breakwaters to reduce aerodynamic forces on FPV systems. Breakwaters, typically used to mitigate wave impacts, can also serve as wind barriers, significantly reducing wind forces before they reach the FPV array. Aerodynamic simulations with and without a breakwater were conducted using CFD to assess this effect. The results show a substantial reduction in lift and drag coefficients, especially for angles of attack up to 10 degrees, demonstrating the effectiveness of the breakwater in protecting the FPV system. However, beyond this threshold, the effectiveness of the breakwater of 2 m reduces. These findings highlight the importance of strategic breakwater placement and heights and their role in enhancing FPV system resilience. The insights gained from this study are critical for optimizing breakwater design and placement, ensuring the structural integrity and performance of FPV systems in varying environmental conditions. The data generated will also contribute to future design improvements for floating PV systems. Full article
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<p>The global FPV concept consisting of breakwater and FPV systems from Sunlit Sea AS.</p>
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<p>Plan of the 420 kWp installation.</p>
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<p>FPV concept from Sunlit Sea AS.</p>
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<p>MABL used in simulations.</p>
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<p>Domain dimension and boundary condition (image taken from ref: [<a href="#B29-energies-17-04873" class="html-bibr">29</a>]).</p>
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<p>Mean pressure coefficients and sectional streamlines at the mid-span (AOA = 30°).</p>
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<p>Mean pressure coefficients and sectional streamlines at the mid-span (AOA = 50°).</p>
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<p>Mean pressure coefficients and sectional streamlines at the mid-span (AOA = 90°).</p>
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<p>Variation of lift force with Number of Panels (Solid lines: CFD results and dashed line: correlation).</p>
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<p>Variation of drag force with Number of Panels (Solid lines: CFD results and dashed line: correlation).</p>
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<p>Contour plots of pressure distribution in the mid-section of 8 × 8 configuration.</p>
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<p>Contour plots of velocity distribution in the mid-section of 8 × 8 configuration.</p>
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<p>Velocity distribution on floats for 8 × 8 configuration.</p>
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<p>A configuration of 8 × 8 FPV with breakwater.</p>
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<p>Velocity distribution on floats for 8 × 8 configuration with breakwater.</p>
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<p>Pressure distribution on floats and breakwater for 8 × 8 configuration with breakwater.</p>
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<p>A configuration of 16 × 16 FPV with breakwater (FPV are not tilted).</p>
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<p>A configuration of 16 × 16 FPV with breakwater (FPV are inclined at 5 degrees).</p>
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17 pages, 102572 KiB  
Article
Improved Lift for Thick Flatback Airfoils in the Inboard Blades of Large Wind Turbines
by Micol Pucci and Stefania Zanforlin
Eng 2024, 5(3), 2345-2361; https://doi.org/10.3390/eng5030122 - 20 Sep 2024
Viewed by 540
Abstract
Thick airfoils are often used in the inboard sections of blades in commercial wind turbines. The main reason for this is to give the blade greater structural strength, but it is well known that thick airfoils degrade aerodynamic performance by stalling at relatively [...] Read more.
Thick airfoils are often used in the inboard sections of blades in commercial wind turbines. The main reason for this is to give the blade greater structural strength, but it is well known that thick airfoils degrade aerodynamic performance by stalling at relatively small angles of attack. The adoption of flatback airfoils instead of sharp trailing edges allows high lift coefficient to be maintained in thick airfoils. In this paper, we propose a novel airfoil design based on a passive flap to further improve the lift coefficient. This new design was tested by numerical simulation on several airfoils with different maximum thickness and different TE thickness. The improved design for flatback airfoils yields a higher lift coefficient, while the drag behaviour is strictly related to the baseline airfoil shape: some airfoils show a decrease in drag at certain angles of attack, while others exhibit a drag increase. In conclusion, the practical implications of the flap’s utilisation on a state-of-the-art blade designed for a 5 MW wind turbine are analysed. The findings demonstrate that, due to the enhanced lift coefficient, it is feasible to shorten the chord while maintaining the power output, thereby reducing material costs. Full article
(This article belongs to the Special Issue Feature Papers in Eng 2024)
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<p>Qualitative representation of a blade element located at radial position <span class="html-italic">r</span>.</p>
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<p>Qualitative representation of the flap proposed in this work. The zoom in the green circle highlights the geometric parameters of the flap (this is merely a qualitative representation; the image is not to scale in order to more clearly depict the flap details).</p>
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<p>Example of a computational grid: (<b>left</b>) full domain extents and (<b>right</b>) detail of the grid near the blade (red rectangle) and near the TE (violet rectangle).</p>
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<p>Example of the calculation grid used for FB airfoils with the flap (<b>left</b>), with the detail of the TE in the red rectangle (<b>right</b>).</p>
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<p>x-y coordinates of the three airfoils used for the validation task.</p>
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<p>(<b>Top</b>) Lift and drag coefficient of the DU97FB and the DU97ST airfoils: comparison between the present 2D simulations and experimental measurements of [<a href="#B14-eng-05-00122" class="html-bibr">14</a>]. (<b>Bottom</b>) Lift and drag coefficient of the FB-3500-1750 airfoil: comparison between the present 2D simulations and experimental measurements of [<a href="#B15-eng-05-00122" class="html-bibr">15</a>].</p>
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<p>Coefficients of lift and drag: comparison between the simple flatback airfoil and the corresponding airfoil with the flap. DU97FB (<b>Top</b>) and FB-3500-1750 (<b>Bottom</b>). In case of DU97FB, the use of the Gurney flap (GF) is also analysed.</p>
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<p>Pressure coefficient at <math display="inline"><semantics> <mrow> <mi>α</mi> <mo>=</mo> <mn>0</mn> <mo>°</mo> </mrow> </semantics></math> and <math display="inline"><semantics> <mrow> <mi>α</mi> <mo>=</mo> <mn>12</mn> <mo>°</mo> </mrow> </semantics></math> (<b>Top</b>) DU97FB with and without the flap (<b>Bottom</b>) FB-3500-1750 with and without the flap.</p>
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<p>Pathlines coloured by static pressure at <math display="inline"><semantics> <mrow> <mi>α</mi> <mo>=</mo> <mn>0</mn> <mo>°</mo> </mrow> </semantics></math> and <math display="inline"><semantics> <mrow> <mi>α</mi> <mo>=</mo> <mn>12</mn> <mo>°</mo> </mrow> </semantics></math> for DU97FB with and without the flap.</p>
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<p>Pathlines coloured by static pressure at <math display="inline"><semantics> <mrow> <mi>α</mi> <mo>=</mo> <mn>0</mn> <mo>°</mo> </mrow> </semantics></math> and <math display="inline"><semantics> <mrow> <mi>α</mi> <mo>=</mo> <mn>12</mn> <mo>°</mo> </mrow> </semantics></math> for FB-3500-1750 with and without the flap.</p>
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<p>RMSE of the mean velocity magnitude at <math display="inline"><semantics> <mrow> <mi>α</mi> <mo>=</mo> <mn>0</mn> <mo>°</mo> </mrow> </semantics></math> and <math display="inline"><semantics> <mrow> <mi>α</mi> <mo>=</mo> <mn>12</mn> <mo>°</mo> </mrow> </semantics></math> for DU97FB with and without the flap.</p>
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<p>RMSE of the mean velocity magnitude at <math display="inline"><semantics> <mrow> <mi>α</mi> <mo>=</mo> <mn>0</mn> <mo>°</mo> </mrow> </semantics></math> and <math display="inline"><semantics> <mrow> <mi>α</mi> <mo>=</mo> <mn>12</mn> <mo>°</mo> </mrow> </semantics></math> for FB-3500-1750 with and without the flap.</p>
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<p>Streamlines coloured by static pressure of the DU97FB airfoil (<b>a</b>), DU97FB airfoil with a flap (<b>b</b>), and DU97FB airfoil with a Gurney flap (<b>c</b>) at <math display="inline"><semantics> <mrow> <mi>α</mi> <mo>=</mo> <mn>10</mn> <mo>°</mo> </mrow> </semantics></math>.</p>
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<p>Enlarged calculation domain for sensitivity analysis. The new large domain includes the old domain, indicated by the blue line in the figure. The domain around the airfoil remains unchanged (red rectangle).</p>
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<p>Coefficients of lift and drag: comparison between the “DU97FB” results (obtained with the <math display="inline"><semantics> <mrow> <mn>28</mn> <mi>c</mi> </mrow> </semantics></math> large domain) and the “DU97FB larger domain” results (obtained with the <math display="inline"><semantics> <mrow> <mn>56</mn> <mi>c</mi> </mrow> </semantics></math> large domain).</p>
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<p>Coefficients of lift and drag: comparison between the simple flatback airfoil and the corresponding airfoil with the flap. DU40FB (<b>Top</b>) and DU35FB (<b>Bottom</b>).</p>
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<p>(<b>Left</b>) Chord distributions with zoom (red box) on the blade portion of interest. (<b>Right</b>) Chord distribution and <math display="inline"><semantics> <mi>ϕ</mi> </semantics></math>, <math display="inline"><semantics> <mi>β</mi> </semantics></math> and <math display="inline"><semantics> <mi>α</mi> </semantics></math> distribution.</p>
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<p>(<b>Left</b>) Torque coefficient along the inboard region of the blade and (<b>Right</b>) thrust coefficient.</p>
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<p>(<b>Left</b>) Torque along the inboard region of the blade and (<b>Right</b>) thrust.</p>
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26 pages, 4772 KiB  
Article
Modelling of Reliability Indicators of a Mining Plant
by Boris V. Malozyomov, Nikita V. Martyushev, Nikita V. Babyr, Alexander V. Pogrebnoy, Egor A. Efremenkov, Denis V. Valuev and Aleksandr E. Boltrushevich
Mathematics 2024, 12(18), 2842; https://doi.org/10.3390/math12182842 - 12 Sep 2024
Viewed by 555
Abstract
The evaluation and prediction of reliability and testability of mining machinery and equipment are crucial, as advancements in mining technology have increased the importance of ensuring the safety of both the technological process and human life. This study focuses on developing a reliability [...] Read more.
The evaluation and prediction of reliability and testability of mining machinery and equipment are crucial, as advancements in mining technology have increased the importance of ensuring the safety of both the technological process and human life. This study focuses on developing a reliability model to analyze the controllability of mining equipment. The model, which examines the reliability of a mine cargo-passenger hoist, utilizes statistical methods to assess failures and diagnostic controlled parameters. It is represented as a transition graph and is supported by a system of equations. This model enables the estimation of the reliability of equipment components and the equipment as a whole through a diagnostic system designed for monitoring and controlling mining equipment. A mathematical and logical model is proposed to calculate availability and downtime coefficients for different structures within the mining equipment system. This analysis considers the probability of failure-free operation of the lifting unit based on the structural scheme, with additional redundancy for elements with lower reliability. The availability factor of the equipment for monitoring and controlling the mine hoisting plant is studied for various placements of diagnostic systems. Additionally, a logistic concept is introduced for organizing preventive maintenance systems and reducing equipment recovery time by optimizing spare parts, integrating them into strategies aimed at enhancing the reliability of mine hoisting plants. Full article
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<p>Generalized block diagram of the measuring system of the diagnostic complex.</p>
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<p>Structural diagram of the measuring unit. The measuring unit is based on a single-chip Intel 8051 microcomputer controlled by the firmware stored in the ROM chip.</p>
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<p>Graph of the mine hoisting machine monitoring and control installation.</p>
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<p>Dependence of the availability factor for monitoring and control of the mine hoisting plant for different ways of diagnostics systems placement: 1—Kg9(F); 2—Kg15(F); 3—Kg12(F); 4—Kg6(F).</p>
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<p>Initial scheme of the mine hoisting plant (MHP) system.</p>
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<p>Transformed scheme.</p>
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<p>Transformations of the bridge circuit: (<b>a</b>)—with absolutely reliable element C, (<b>b</b>)—failed element C.</p>
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<p>Variation of probability of failure-free operation of a mine hoisting unit: <span class="html-italic">P</span>—probability of failure-free operation of the initial scheme; <span class="html-italic">P</span>′—probability of failure-free operation of the scheme with increasing reliability of individual elements; <span class="html-italic">P</span>″—probability of failure-free operation of the scheme with application of structural redundancy.</p>
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<p>Dependence of the probability of failure-free operation of the “2 out of 4” system on the probability of failure-free operation of its elements.</p>
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<p>Structural diagram of the system after structural redundancy.</p>
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<p>Graph of dependence of the guaranteed probability of ensuring the stock of lifting elements on the parameter <math display="inline"><semantics> <mrow> <msub> <mi>z</mi> <mrow> <mi>av</mi> </mrow> </msub> </mrow> </semantics></math> at different values of <math display="inline"><semantics> <mrow> <msub> <mi>N</mi> <mrow> <mi>s</mi> <mi>p</mi> <mi>e</mi> </mrow> </msub> </mrow> </semantics></math>. The numbers above curves 1,2,....10 denote the values of the parameter <span class="html-italic">N<sub>spe</sub></span> used to plot these curves.</p>
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<p>Graph of dependence of the stock factor of the lifting unit elements on the average number of their failures <math display="inline"><semantics> <mrow> <msub> <mi>z</mi> <mrow> <mi>av</mi> </mrow> </msub> </mrow> </semantics></math> for the overhaul period. Curve 1—probability of providing the reserve <span class="html-italic">P<sub>pr</sub></span> = 0.95; curve 2—probability of providing the reserve <span class="html-italic">P<sub>pr</sub></span> = 0.9.</p>
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18 pages, 5315 KiB  
Article
Investigation of a Tube-Launched Unmanned Aerial Vehicle with a Variable-Sweep Wing
by Peng Si, Mingjian Wu, Yongqing Huo and Zhilin Wu
Drones 2024, 8(9), 474; https://doi.org/10.3390/drones8090474 - 10 Sep 2024
Viewed by 838
Abstract
Foldable wings are designed for tube-launched unmanned aerial vehicles (UAVs), aiming to improve portability and meet launch platform requirements. However, conventional tube-launched UAVs cannot operate across the wide speed ranges required for the performance of multiple missions, due to the fixed configuration of [...] Read more.
Foldable wings are designed for tube-launched unmanned aerial vehicles (UAVs), aiming to improve portability and meet launch platform requirements. However, conventional tube-launched UAVs cannot operate across the wide speed ranges required for the performance of multiple missions, due to the fixed configuration of their wings after launch. This study therefore proposes a tube-launched UAV which can change wing-sweep angle to expand the flight speed range and enhance the UAV’s agility. A computational aerodynamics method is employed to assess the transient aerodynamic performance of the UAV during the sweep morphing process. The simulation results indicate that the transient aerodynamic forces generate a dynamic hysteresis loop around the quasi-steady data. The lift and drag coefficients exhibit maximum relative deviations of 18.5% and 12.7% from the quasi-steady data for the sweep morphing period of 0.5 s. The hysteresis effect of the flow structure, rather than the additional velocity resulting from wing-sweep morphing, is the major contributor to the aerodynamic hysteresis loop. Compared to the conventional tube-launched UAVs, the proposed tube-launched UAV with a variable-sweep wing shows a wider flight speed range, from 22.59 to 90.12 m/s, and achieves an 82.84% increase in loitering speed. To verify the effectiveness of the wing-sweeping concept, a prototype was developed, and a flight test was carried out. The test data obtained from flight control system agree well with the simulation data, which demonstrates the feasibility and effectiveness of the variable-sweep wing in widening the speed range for tube-launched UAVs. This work can provide a reference for the design of tube-launched UAVs for wide speed range flight. Full article
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<p>Prototype of the tube-launched UAV with a variable-sweep wing: (<b>a</b>) loitering configuration; (<b>b</b>) high-speed configuration; (<b>c</b>) dash configuration; and (<b>d</b>) folded configuration.</p>
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<p>Mesh details of the tube-launched UAV with a variable-sweep wing: (<b>a</b>) assembly mesh; (<b>b</b>) detailed mesh near the wing root; and (<b>c</b>) mesh of the folded configuration.</p>
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<p>Fluid domain with boundary conditions.</p>
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<p>Validation of mesh size: (<b>a</b>) lift coefficient and (<b>b</b>) drag coefficient.</p>
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<p>Validation of transient time step size: (<b>a</b>) lift coefficient and (<b>b</b>) drag coefficient.</p>
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<p>Transient aerodynamic coefficients during the sweep morphing process under four morphing periods: (<b>a</b>) lift coefficient and (<b>b</b>) drag coefficient.</p>
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<p>Relative deviations of the transient aerodynamic coefficients from the corresponding quasi-steady data throughout the sweep-backward process: (<b>a</b>) lift coefficient and (<b>b</b>) drag coefficient.</p>
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<p>Aerodynamic hysteresis loops of the wing, fuselage, and tail during the sweep morphing at <span class="html-italic">T</span> = 0.5 s.</p>
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<p>Aerodynamic hysteresis loops, including and excluding the wing velocity, in the numerical simulations.</p>
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<p>Comparisons of pressure coefficient on wing surface excluding and including the wing velocity from the wing-sweep motion.</p>
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<p>Comparisons of shear stress on the wing surface, excluding and including the wing velocity from the wing-sweep motion: (<b>a</b>) upper surface and (<b>b</b>) lower surface.</p>
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<p>Pressure contour for the whole aircraft under three morphing states (<span class="html-italic">T</span> = 0.5 s, <span class="html-italic">δ</span> = 30°): (<b>a</b>) sweep forward; (<b>b</b>) sweep backward; and (<b>c</b>) quasi-steady.</p>
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<p>Pressure coefficient distribution on the cross-section of the section plane and the left wing: (<b>a</b>) pressure coefficient distribution; and (<b>b</b>) slice plane.</p>
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<p>Vorticity contour at the section plane near the combination of the wing and the tube: (<b>a</b>) sweep forward; (<b>b</b>) sweep backward; (<b>c</b>) quasi-steady; and (<b>d</b>) the slice plane.</p>
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<p>Velocity contours at a spanwise section under three morphing states (<span class="html-italic">T</span> = 0.5 s, <span class="html-italic">y</span>/<span class="html-italic">b</span> = 0.5, <span class="html-italic">δ</span> = 30°): (<b>a</b>) sweep forward; (<b>b</b>) sweep backward; and (<b>c</b>) quasi-steady.</p>
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<p>Pressure contours at a spanwise section under three morphing states (<span class="html-italic">T</span> = 0.5 s, <span class="html-italic">y</span>/<span class="html-italic">b</span> = 0.5, <span class="html-italic">δ</span> = 30°): (<b>a</b>) sweep forward, (<b>b</b>) sweep backward, and (<b>c</b>) quasi-steady.</p>
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<p>Pressure coefficient distributions under three morphing states (<span class="html-italic">T</span> = 0.5 s, <span class="html-italic">y</span>/<span class="html-italic">b</span> = 0.5, <span class="html-italic">δ</span> = 30°).</p>
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<p>Transient aerodynamic moment coefficient and location of the center of pressure versus the sweep angle during the sweep morphing process under four morphing periods: (<b>a</b>) location of the center of pressure and (<b>b</b>) nose-down pitching moment coefficient.</p>
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<p>Flight speed of the tube-launched UAV versus angle of attack under three different wing-sweep angles.</p>
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<p>Flight performance of the tube-launched UAV under different wing-sweep angles at an angle of attack of 4°: (<b>a</b>) loitering speed and response time; and (<b>b</b>) power consumption and flight endurance.</p>
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<p>Flight testing of the tube-launched aircraft, morphing into the loiter configuration.</p>
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<p>Flight testing of the tube-launched aircraft, morphing into the high-speed configuration.</p>
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15 pages, 12994 KiB  
Article
Numerical Analysis of Leading-Edge Roughness Effects on the Aerodynamic Performance of a Thick Wind Turbine Airfoil
by Wei Zhang, Kuichao Ma, Chang Cai, Xiangyu Sun, Jun Zhang, Xiaohui Zhong, Xiaomin Rong and Qing’an Li
J. Mar. Sci. Eng. 2024, 12(9), 1588; https://doi.org/10.3390/jmse12091588 - 8 Sep 2024
Viewed by 548
Abstract
The aerodynamic performance of wind turbine airfoils is crucial for the efficiency and reliability of wind energy systems, with leading-edge roughness significantly impacting blade performance. This study conducts numerical simulations on the DU 00-W-401 airfoil to investigate the effects of leading-edge roughness. Results [...] Read more.
The aerodynamic performance of wind turbine airfoils is crucial for the efficiency and reliability of wind energy systems, with leading-edge roughness significantly impacting blade performance. This study conducts numerical simulations on the DU 00-W-401 airfoil to investigate the effects of leading-edge roughness. Results reveal that the rough airfoil exhibits a distinctive “N”-shaped lift coefficient curve. The formation mechanism of this nonlinear lift curve is primarily attributed to the development of the trailing-edge separation vortex and variations in the adverse pressure gradient from the maximum thickness position to the trailing-edge confluence. The impact of different roughness heights is further investigated. It is discovered that when the roughness height is higher than 0.3 mm, the boundary layer can be considered fully turbulent, and the lift curve shows the “N” shape stably. When the roughness height is between 0.07 mm and 0.1 mm, a transitional state can be observed, with several saltation points in the lift curve. The main characteristics of different flow regimes based on different lift curve segments are summarized. This research enhances the understanding of the effects of leading-edge roughness on the aerodynamic performance of a thick wind turbine airfoil, and the simulation method for considering the effect of leading-edge roughness is practical to be applied on large-scale wind turbine blade to estimate the aerodynamic performance under rough leading-edge conditions, thereby supporting advancements in wind turbine technology and promoting the broader adoption of renewable energy. Full article
(This article belongs to the Special Issue Advances in Offshore Wind—2nd Edition)
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<p>Profile of DU 00-W-401 airfoil and the position of leading-edge roughness.</p>
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<p>Mesh and boundary condition.</p>
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<p>Convergence history of the lift and drag coefficient at 14° AOA.</p>
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<p>Lift coefficient of the smooth and rough airfoils (experimental data is adapted from Ref. [<a href="#B17-jmse-12-01588" class="html-bibr">17</a>]).</p>
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<p>Drag coefficient of the smooth and rough airfoils.</p>
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<p>Contour of the turbulence intensity around the smooth airfoil at 0° AOA.</p>
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<p>Nondimensional streamwise wall shear around the smooth airfoil at 0° AOA.</p>
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<p>Contour of the turbulence intensity around the rough airfoil at 0° AOA.</p>
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<p>Nondimensional streamwise wall shear around the rough airfoil at 0° AOA.</p>
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<p>Streamlines around the smooth airfoil at different AOAs.</p>
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<p>Streamlines around the rough airfoil at different AOAs.</p>
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<p>Pressure contour and streamlines around the rough airfoil in the negative-slope lift region. The green and purple arrows represent the characteristic streamlines between the main stream and the separation vortex.</p>
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<p>Pressure coefficient distribution on the rough airfoil in the negative-slope lift region.</p>
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<p>Lift coefficient of the rough airfoil with different roughness heights.</p>
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<p>Definition of different flow regimes based on different lift curve segments.</p>
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<p>Influence of roughness height on different flow regimes.</p>
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<p>Nondimensional streamwise wall shear around the rough airfoil at 10° AOA (RH = 0.10 mm and 0.30 mm).</p>
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19 pages, 5637 KiB  
Article
A Zonal Detached Eddy Simulation of the Trailing Edge Stall Process of a LS0417 Airfoil
by Wenbo Shi, Heng Zhang and Yuanxiang Li
Aerospace 2024, 11(9), 731; https://doi.org/10.3390/aerospace11090731 - 6 Sep 2024
Viewed by 438
Abstract
A Zonal Detached Eddy Simulation (ZDES) based on the SST turbulence model is implemented to the numerical investigation of the trailing edge stall of a LS-0417 airfoil, which includes multiple DES modes for different classifications of flow separation and adopts the subgrid scale [...] Read more.
A Zonal Detached Eddy Simulation (ZDES) based on the SST turbulence model is implemented to the numerical investigation of the trailing edge stall of a LS-0417 airfoil, which includes multiple DES modes for different classifications of flow separation and adopts the subgrid scale definition of Δω. The entire stall process under a series of AOA is simulated according to the experiment condition. The performance of URANS and ZDES in the prediction of the stall flow field are compared. The results reveal that the stall point obtained through ZDES is consistent with the experiment; the deviation of the predicted maximum lift coefficient from the measured result is only 0.8%, while the maximum lift is overpredicted by both RANS and URANS. The high frequency fluctuations are observed in the time history of the lift in ZDES result during stall. With the increase in the AOA, a mild development of separation and a gradual decrease in leading edge peak suction are manifested in the ZDES result. The alternate shedding of shear layers and the interference between the leading edge and trailing edge vortices are illustrated through ZDES near the stall point; the corresponding turbulent fluctuations with high intensity are captured in the separation region, which indicates the essential difference in the prediction of stall process between URANS and ZDES. Full article
(This article belongs to the Section Aeronautics)
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<p>The definition of <math display="inline"><semantics> <mrow> <msub> <mo>Δ</mo> <mi>ω</mi> </msub> </mrow> </semantics></math> of a grid cell.</p>
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<p>The topology and section of the computational grid of LS-0417.</p>
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<p>The definition of calculation region of ZDES.</p>
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<p>The convergence history of the lift coefficient with different methods.</p>
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<p>The convergence history of lift coefficient with different time steps.</p>
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<p>Comparison of lift prediction results with different methods.</p>
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<p>The convergence history of lift coefficient obtained with URANS and ZDES at different AOAs.</p>
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<p>The convergence history of lift coefficient obtained with URANS and ZDES at different AOAs.</p>
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<p>Comparison of time-averaged streamlines between URANS and ZDES.</p>
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<p>Comparison of time-averaged streamlines between URANS and ZDES.</p>
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<p>Comparison of time-averaged pressure distribution between URANS and ZDES.</p>
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<p>Comparison of instantaneous vorticity between URANS and ZDES.</p>
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<p>Comparison of instantaneous vorticity between URANS and ZDES.</p>
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<p>Comparison of time-averaged turbulent fluctuation between URANS and ZDES, α = 20°.</p>
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<p>Comparison of iso-surface of Q-criterion between URANS and ZDES.</p>
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16 pages, 6066 KiB  
Article
Effect of Gurney Flaps on Non-Planar Wings at Low Reynolds Number
by Lance W. Traub
Aerospace 2024, 11(9), 728; https://doi.org/10.3390/aerospace11090728 - 6 Sep 2024
Viewed by 375
Abstract
The effect of spanwise wing non-planarity, employed in conjunction with a Gurney flap, is presented. Testing was undertaken in a low-speed wind tunnel using a rectangular wing with an aspect ratio of three. The outer one-third of the wing was non-planar, which took [...] Read more.
The effect of spanwise wing non-planarity, employed in conjunction with a Gurney flap, is presented. Testing was undertaken in a low-speed wind tunnel using a rectangular wing with an aspect ratio of three. The outer one-third of the wing was non-planar, which took the form of either dihedral or a circular arc. A 2% high Gurney flap was implemented such that it could extend over the entire span or the planar inboard section. The loads were measured using a sting balance. The data show that non-planarity increases the maximum lift coefficient and the wing’s lift curve slope. Gurney flap lift modulation was enhanced in the presence of non-planarity. The addition of Gurney flaps caused a greater increment in the minimum drag coefficient for the non-planar wings. The Gurney flaps reduced the lift-dependent drag of the wings. As a whole, the Gurney flaps reduced the maximum lift-to-drag ratio (L/D)max for the non-planar wings; however, the flat wing exhibited a small L/D increment with flap addition. Full article
(This article belongs to the Section Aeronautics)
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<p>Model geometry, all dimensions in inches.</p>
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<p>Effect of GF and non-planarity on CL, data grouped by (<b>a</b>) outboard section type and (<b>b</b>) flap distribution.</p>
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<p>Effect of the spanwise distribution of the flap on the lift increment caused by the GF relative to the wing without a flap.</p>
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<p>Comparison of experimentally measured C<sub>L</sub> data with analytical predictions for different outboard wing sections.</p>
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<p>Effect of the GF and non-planarity on the measured (<b>a</b>) drag polar and (<b>b</b>) linearized drag polar, with data grouped by outboard section geometry. (<b>c</b>) Surface skin friction lines indicating the location of the laminar separation bubble.</p>
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<p>Effect of the GF and non-planarity on the measured (<b>a</b>) drag polar and (<b>b</b>) linearized drag polar, with data grouped by outboard section geometry. (<b>c</b>) Surface skin friction lines indicating the location of the laminar separation bubble.</p>
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<p>Effect of the GF and non-planarity on the measured (<b>a</b>) drag polar and (<b>b</b>) linearized drag polar, with the data grouped by GF distribution.</p>
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<p>Drag polar comparison with AVL predictions as affected by outboard wing geometry.</p>
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<p>Comparison of a flat wing with equivalent arc length to the non-planar wing; effect on the drag polar.</p>
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<p>Effect of the GF and non-planarity on the lift-to-drag ratio, with data grouped by (<b>a</b>) outboard section type and (<b>b</b>) spanwise distribution of the flap.</p>
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<p>Effect of non-planarity and outboard section geometry on the benefit margin provided by the Gurney flap.</p>
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<p>Effect of the GF and non-planarity on C<sub>m</sub>, with data grouped by (<b>a</b>) outboard section type and (<b>b</b>) spanwise distribution of the flap.</p>
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25 pages, 6088 KiB  
Article
Production Prediction and Influencing Factors Analysis of Horizontal Well Plunger Gas Lift Based on Interpretable Machine Learning
by Jinbo Liu, Haowen Shi, Jiangling Hong, Shengyuan Wang, Yingqiang Yang, Honglei Liu, Jiaojiao Guo, Zelin Liu and Ruiquan Liao
Processes 2024, 12(9), 1888; https://doi.org/10.3390/pr12091888 - 3 Sep 2024
Viewed by 698
Abstract
With the development of unconventional natural gas resources, plunger gas lift technology has gained widespread application. Accurately predicting gas production from unconventional gas reservoirs is a crucial step in evaluating the effectiveness of plunger gas lift technology and optimizing its design. However, most [...] Read more.
With the development of unconventional natural gas resources, plunger gas lift technology has gained widespread application. Accurately predicting gas production from unconventional gas reservoirs is a crucial step in evaluating the effectiveness of plunger gas lift technology and optimizing its design. However, most existing prediction methods are mechanism-driven, incorporating numerous assumptions and simplifications that make it challenging to fully capture the complex physical processes involved in plunger gas lift technology, ultimately leading to significant errors in capacity prediction. Furthermore, engineering design factors and production system factors associated with plunger gas lift technology can contribute to substantial deviations in gas production forecasts. This study employs three powerful regression algorithms, XGBoost, Random Forest, and SVR, to predict gas production in plunger gas lift wells. This method comprehensively leverages various types of data, including collected engineering design, production system, and production data, directly extracting the underlying patterns within the data through machine learning algorithms to establish a prediction model for gas production in plunger gas lift wells. Among these, the XGBoost algorithm stands out due to its robustness and numerous advantages, such as high accuracy, ability to effectively handle outliers, and reduced risk of overfitting. The results indicate that the XGBoost algorithm exhibits impressive performance, achieving an R2 (coefficient of determination) value of 0.87 for six-fold cross-validation and 0.85 for the test set. Furthermore, to address the “black box” problem (the inability to know the internal working structure and workings of the model and to directly understand the decision-making process), which is commonly associated with conventional machine learning models, the SHAP (Shapley additive explanations) method was utilized to globally and locally interpret the established machine learning model, analyze the main factors (such as starting time of wells, gas–liquid ratio, catcher well inclination angle, etc.) influencing gas production, and enhance the credibility and transparency of the model. Taking plunger gas lift wells in southwest China as an example, the effectiveness and practicality of this method are demonstrated, providing reliable data support for shale gas production prediction, and offering valuable guidance for actual on-site production. Full article
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<p>Workflow diagram.</p>
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<p>Entropy diagram.</p>
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<p>Mutual information diagram.</p>
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<p>Sixfold cross-validation schematic diagram.</p>
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<p>Proportion of missing data values.</p>
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<p>Histogram of characteristic data and gas production.</p>
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<p>Histogram of characteristic data and gas production.</p>
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<p>Pearson correlation heatmap.</p>
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<p>Mutual information values of 11 gas production values relative to plunger gas lift wells.</p>
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<p>Prediction graph and error statistics graph of XGBoost model.</p>
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<p>Prediction graph and error statistics graph of random forest model.</p>
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<p>Prediction graph and error statistics graph of SVR model.</p>
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<p>XGBoost model SHAP value global interpretation.</p>
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<p>XGBoost model SHAP value local interpretation.</p>
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<p>Global interpretation of SHAP values of Random Forest model.</p>
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<p>Local interpretation of SHAP values of random forest model.</p>
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21 pages, 7151 KiB  
Article
Establishment and Research of Cotton Stalk Moisture Content–Discrete Element Parameter Model Based on Multiple Verification
by Tao Wu, Limin Yan, Deli Jiang, Haixiao Gou, Xuanhe Fu and Jinhao Zhang
Processes 2024, 12(8), 1770; https://doi.org/10.3390/pr12081770 - 21 Aug 2024
Viewed by 519
Abstract
In view of the large difference in moisture content of cotton stalk in autumn in Xinjiang, the existing process of obtaining discrete element simulation parameters of cotton stalk is low in accuracy and complicated in operation, leading to the problems of poor universality [...] Read more.
In view of the large difference in moisture content of cotton stalk in autumn in Xinjiang, the existing process of obtaining discrete element simulation parameters of cotton stalk is low in accuracy and complicated in operation, leading to the problems of poor universality and low accuracy in regard to the discrete element simulation parameter-calibration method in the process of mechanized transportation, throwing and returning to the field. Therefore, the experimental study on cotton stalk with different moisture content was carried out with the accumulation angle as the response value, so as to construct a parameter model that can quickly and accurately calibrate cotton stalk with different levels of moisture content. The model has high applicability and flexibility, and it can be widely used in the simulation test of various cotton field-operation machinery, such as a residual film-recycling machine, cotton picker, crushing and returning machine and other equipment. The water content–accumulation angle model was established by the cylinder-lifting method, and the correlation coefficient of the model was 0.9993. Based on EDEM 2020 software, the Hertz–Mindlin model was used to simulate the stacking angle of cotton stalk, and the rolling friction coefficient, static friction coefficient and collision recovery coefficient between cotton stalk and cotton stalk–steel were obtained. Through the Plackett–Burman test, climbing test and Box–Behnken test, three significant parameters, namely the rolling friction coefficient, static friction coefficient and static friction coefficient between cotton stalk and steel, were selected from discrete element simulation parameters to characterize the moisture content of cotton stalk, and the accumulation angle–discrete element parameter model was established. The p-value of the model was less than 0.0001, and the relative error was only 2.67%. Based on the moisture content–stacking angle model and the stacking angle–discrete element parameter model, the moisture content–discrete element parameter model was constructed. The model was verified by the cylinder-lifting method and the plate-drawing method, and the relative error was only 2.79%. Finally, the model was further verified by comparing the effect of the throwing uniformity between the mechanical simulation test and field test, and the relative error was only 4.75%. The test proves that the moisture content–discrete element parameter model is accurate and reliable, not only providing the design basis and support for the mechanization research of cotton stalk conveying and returning to the field in Xinjiang but also providing ideas for the calibration of discrete element simulation parameters of other crop straws. Full article
(This article belongs to the Section Process Control and Monitoring)
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<p>Change in cotton stalk moisture content from 8 October to 28 October.</p>
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<p>Cylinder-lifting method’s accumulation angle-measuring instrument.</p>
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<p>Measurement of cotton stalk stacking angle.</p>
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<p>Measurement of cotton stalk stacking angle.</p>
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<p>Geometric model of cotton stalk particles.</p>
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<p>Cylinder-lifting method’s simulation test.</p>
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<p>Image processing.</p>
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<p>Accumulation angle simulation and physical experiment based on plate-pulling method.</p>
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<p>Dispersing and returning field-testing machine.</p>
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<p>Fitting curve of moisture content–stacking angle of cotton stalk.</p>
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<p>Simulation and physical test stacking angle.</p>
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<p>Simulation and physical test stacking angle.</p>
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<p>Physical test and simulation test of plate-drawing method.</p>
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<p>Physical test and simulation test of cylinder-lifting method.</p>
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<p>The effect and statistical method of experimental scattering being returned to the field.</p>
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19 pages, 7542 KiB  
Article
A Method to Design an Efficient Airfoil for Small Wind Turbines in Low Wind Speed Conditions Using XFLR5 and CFD Simulations
by Le Quang Sang, Tinnapob Phengpom, Dinh Van Thin, Nguyen Huu Duc, Le Thi Thuy Hang, Cu Thi Thanh Huyen, Nguyen Thi Thu Huong and Quynh T. Tran
Energies 2024, 17(16), 4113; https://doi.org/10.3390/en17164113 - 19 Aug 2024
Viewed by 725
Abstract
Small wind turbines operating in low wind speed regions have not had any significant success. In addition, small wind speed regions occupy a large area of the world, so they represent a potential area for installing small wind turbines in the future. In [...] Read more.
Small wind turbines operating in low wind speed regions have not had any significant success. In addition, small wind speed regions occupy a large area of the world, so they represent a potential area for installing small wind turbines in the future. In this paper, a method to design an efficient airfoil for small wind turbines in low wind speed conditions using XFLR5 and CFD simulations is implemented. Because the impact of the airflow on the blade surface under low Re number conditions can change suddenly for small geometries, designing the airfoil shape to optimize the aerodynamic performance is essential. The tuning of the key geometric parameters using inversion techniques for better aerodynamic performance is presented in this study. A two-dimensional model was used to consider the airflow on the airfoil surface with differences in the angle of attack. The original S1010 airfoil was used to design a new airfoil for increasing the aerodynamic efficiency by using V6.57 XFLR5 software. Subsequently, the new VAST-EPU-S1010 airfoil model was adjusted to the maximum thickness and the maximum thickness position. It was simulated in low wind speed conditions of 4–6 m/s by a computational fluid dynamics simulation. The lift coefficient, drag coefficient, and CL/CD coefficient ratio were evaluated under the effect of the angle of attack and the maximum thickness by using the k-ε model. The simulation results show that the VAST-EPU-S1010 airfoil achieved the greatest aerodynamic efficiency at an angle of attack of 3°, a maximum thickness of 8%, and a maximum thickness position of 20.32%. The maximum value of CL/CD of the new airfoil at 6 m/s was higher than at 4 m/s by about 6.25%. Full article
(This article belongs to the Section A3: Wind, Wave and Tidal Energy)
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<p>Two-dimensional design of S1010 airfoil.</p>
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<p><span class="html-italic">C<sub>L</sub></span> and <span class="html-italic">C<sub>D</sub></span> coefficients of NACA64A010 according to the angle of attack: (<b>a</b>) <span class="html-italic">C<sub>L</sub></span> coefficient; and (<b>b</b>) <span class="html-italic">C<sub>D</sub></span> coefficient.</p>
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<p><span class="html-italic">C<sub>L</sub></span> and <span class="html-italic">C<sub>D</sub></span> coefficients of S1210 according to the angle of attack: (<b>a</b>) <span class="html-italic">C<sub>L</sub></span> coefficient; and (<b>b</b>) <span class="html-italic">C<sub>D</sub></span> coefficient.</p>
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<p><span class="html-italic">C<sub>L</sub></span>/<span class="html-italic">C<sub>D</sub></span> coefficient of S1010 changes with AoA at wind speed of 5 m/s.</p>
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<p>The (<span class="html-italic">C<sub>L</sub></span>/<span class="html-italic">C<sub>D</sub></span>)<sub>max</sub> coefficient depends on the different airfoil models.</p>
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<p>Three-dimensional design of VAST-EPU-S1010 airfoil.</p>
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<p>Computational domain around VAST-EPU-S1010 airfoil.</p>
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<p>Mesh around VAST-EPU-S1010 airfoil.</p>
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<p>Effect of AoA on <span class="html-italic">C<sub>L</sub></span> and <span class="html-italic">C<sub>D</sub></span> coefficients and <span class="html-italic">C<sub>L</sub></span>/<span class="html-italic">C<sub>D</sub></span> ratio at various speeds.</p>
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<p>Effect of the wind speed on the maximum <span class="html-italic">C<sub>L</sub></span>/<span class="html-italic">C<sub>D</sub></span> ratio values.</p>
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<p>Velocity contour comparative analysis of the impact of the angle of attack (AoA) at various speeds.</p>
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<p>Velocity vector comparative analysis of the impact of the angle of attack (AoA) at various speeds.</p>
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<p>Optimal value of <span class="html-italic">C<sub>L</sub></span>/<span class="html-italic">C<sub>D</sub></span> ratio at the AoA = 3°, according to different maximum thicknesses.</p>
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<p>Velocity contour with thicknesses varying from 6–9%: (<b>a</b>) t/c = 6%; (<b>b</b>) t/c = 7%; (<b>c</b>) t/c = 8%; and (<b>d</b>) t/c = 9%.</p>
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