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Search Results (1,741)

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15 pages, 3147 KiB  
Article
Transmission Line Icing Prediction Based on Physically Guided Fast-Slow Transformer
by Feng Wang and Ziming Ma
Energies 2025, 18(3), 695; https://doi.org/10.3390/en18030695 - 3 Feb 2025
Viewed by 105
Abstract
To improve the accuracy of the icing prediction model for overhead transmission lines, a physics-guided Fast-Slow Transformer icing prediction model for overhead transmission lines is proposed, which is based on the icing prediction model with meteorological input characteristics. First, the ice cover data [...] Read more.
To improve the accuracy of the icing prediction model for overhead transmission lines, a physics-guided Fast-Slow Transformer icing prediction model for overhead transmission lines is proposed, which is based on the icing prediction model with meteorological input characteristics. First, the ice cover data is segmented into different time resolutions through Fourier transform; a transformer model based on Fourier transform is constructed to capture the local and global correlations of the ice cover data; then, according to the calculation model of the comprehensive load on the conductor and the conductor state equation, the variation law of ice thickness, temperature, wind speed, and tension is analyzed, and the model loss function is constructed according to the variation law to guide the training process of the model. Finally, the sample mixing enhancement algorithm is used to reduce the overfitting problem and improve the generalization performance of the prediction model. The results show that the proposed prediction model can consider the mechanical constraints in the ice growth process and accurately capture the dependence between ice cover and meteorology. Compared with traditional prediction models such as LSTM (Long Short-Term Memory) networks, its mean square error, mean absolute error, and mean absolute percentage error are reduced by 0.464–0.674, 0.41–0.53, and 8.87–11.5%, respectively, while the coefficient of determination (R2) is increased by 0.2–0.29. Full article
(This article belongs to the Section F5: Artificial Intelligence and Smart Energy)
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<p>Transmission line stress model.</p>
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<p>Loss function structure diagram.</p>
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<p>Prediction flowchart.</p>
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<p>FSFormer structure.</p>
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<p>Comparison of prediction results.</p>
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<p>Prediction error box plot.</p>
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<p>Model physical inconsistency.</p>
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<p>Comparison of performance between different <span class="html-italic">T</span> and <span class="html-italic">k</span>.</p>
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<p>Probability density of Beta distribution at different <span class="html-italic">α</span>.</p>
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<p>Comparison of different α performances.</p>
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12 pages, 960 KiB  
Article
Motivation for Participation in Civil Wind Bands: Contributions for Non-Formal Educational Contexts
by José Cidade, Alexandra Sá Costa and João Caramelo
Educ. Sci. 2025, 15(2), 173; https://doi.org/10.3390/educsci15020173 - 2 Feb 2025
Viewed by 215
Abstract
Portuguese civil wind bands have operated as voluntary, non-profit organisations since the 19th century and serve as presentational and communal platforms for amateur music-making. Their core mission centres on providing music instruction and practical training for amateur musicians. This study examines the motivational [...] Read more.
Portuguese civil wind bands have operated as voluntary, non-profit organisations since the 19th century and serve as presentational and communal platforms for amateur music-making. Their core mission centres on providing music instruction and practical training for amateur musicians. This study examines the motivational factors driving adult musicians’ participation in civil wind bands. The research involved 617 adult wind band musicians nationwide who completed an online questionnaire. The findings indicate that fellowship consistently ranks as the primary motivator for participation, regardless of gender, age, and formal music education level. Musicianship emerged as the second most influential factor, with younger and older musicians placing substantial value on personal musical growth. Conversely, conductor leadership was the least important motivator, particularly among older musicians and those with higher levels of formal music training. These findings highlight the multidimensional nature of motivations for sustained participation in civil wind bands. The implications suggest that music directors and organisational managers can leverage insights from motivational studies to foster inclusive, self-rewarding, and intergenerational participation. Full article
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<p>Kruskal–Wallis’s test results for personal musicianship subscale and significant difference between age groups.</p>
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<p>Kruskal–Wallis’s test results for conductor’s leadership subscale and significant difference between age groups.</p>
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<p>Kruskal–Wallis’s test results for conductor’s leadership subscale and significant differences between music education level groups.</p>
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14 pages, 4418 KiB  
Article
Controlling the All-Solid Surface Reaction Between an Li1.3Al0.3Ti1.7(PO4)3 Electrolyte and Anode Through the Insertion of Ag and Al2O3 Nano-Interfacial Layers
by Gwanhee Song, Bojoong Kim, Inkook Hwang, Jiwon Kim, Jinmo Kim and Chang-Bun Yoon
Materials 2025, 18(3), 609; https://doi.org/10.3390/ma18030609 - 29 Jan 2025
Viewed by 479
Abstract
Solid-state lithium batteries are considered ideal due to the safety of solid-state electrolytes. The Na superionic conductor-type Li1.3Al0.3Ti1.7(PO4)3 (LATP) is a solid electrolyte with high ionic conductivity, low cost, and stability. However, LATP is [...] Read more.
Solid-state lithium batteries are considered ideal due to the safety of solid-state electrolytes. The Na superionic conductor-type Li1.3Al0.3Ti1.7(PO4)3 (LATP) is a solid electrolyte with high ionic conductivity, low cost, and stability. However, LATP is reduced upon contact with metallic lithium, leading to lithium dendrite growth on the anode during charging. In this study, LATP was synthesized, and the relationship between crystallinity and ionic conductivity was investigated at different heat treatment temperatures. Optimal sintering conditions and ionic conductivity were analyzed for sintering temperatures from 800 to 1000 °C. To suppress reactions with Li metal, 50 nm thick Ag and 10 nm thick Al2O3 layers were deposited on LATP via DC sputtering and plasma-enhanced atomic layer deposition. The electrochemical stability was tested under three conditions: uncoated LATP, Al2O3-coated LATP, and Ag+Al2O3-coated LATP. The stability improved in the following order: uncoated < Al2O3-coated < Ag+Al2O3-coated. The Al2O3 coating suppressed secondary phase formation by preventing direct contact between LATP and Li, while Ag coating mitigated charge concentration, inhibiting dendrite growth. These findings demonstrate that Ag and Al2O3 nano-layers enhance electrolyte stability, advancing solid-state battery reliability and commercialization. Full article
(This article belongs to the Special Issue Ionic Liquid Electrolytes for Energy Storage Devices)
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<p>X-ray diffraction patterns of Li<sub>1.3</sub>Al<sub>0.3</sub>Ti<sub>1.7</sub>(PO4)<sub>3</sub> (LATP) at different sintering temperatures (850 °C to 1000 °C).</p>
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<p>Cross-sectional scanning electron microscopy (SEM) images of the LATP solid electrolytes at different temperatures, (<b>a</b>) 850 °C, (<b>b</b>) 900 °C, (<b>c</b>) 950 °C, and (<b>d</b>) 1000 °C, for a sintering time of 6 h.</p>
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<p>Measured absolute and relative densities of LATP as a function of the sintering temperature.</p>
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<p>Ionic conductivity of LATP at different sintering temperatures: (<b>a</b>) conductivity graph; (<b>b</b>) temperature-specific data.</p>
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<p>Activation energy measurements for LATP sintered at 950 °C: (<b>a</b>) ln σ vs. 1000/T plot graph; (<b>b</b>) Nyquist plot results at different temperatures.</p>
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<p>Mechanisms observed in LATP-based all-solid-state batteries during charge–discharge cycles: (<b>a</b>) formation of secondary phases and dendrites, (<b>b</b>) effects of Al<sub>2</sub>O<sub>3</sub> nano-layer coating, and (<b>c</b>) effects of Ag/Al<sub>2</sub>O<sub>3</sub> nano-layer coating.</p>
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<p>Performance evaluation of Li–Li symmetric cells with LATP sintered samples: bare LATP, 10 nm Al<sub>2</sub>O<sub>3</sub>-coated sample, and 10 nm Al<sub>2</sub>O<sub>3</sub>/50 nm Ag-coated sample. (<b>a</b>) Full DC cycling graph. (<b>b</b>) Maximum and minimum voltage change graph. (<b>c</b>) Charge–discharge profile at 0 h (<b>d</b>) Charge–discharge profile at 200 h. (<b>e</b>) Voltage change (Vmax−Vmin) after 200 h for each coating type.</p>
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<p>Cross-sectional analysis after 200 h of DC cycling: (<b>a</b>) LATP only, (<b>b</b>) Al<sub>2</sub>O<sub>3</sub>-coated LATP, and (<b>c</b>) Al<sub>2</sub>O<sub>3</sub>/Ag-coated LATP.</p>
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16 pages, 2646 KiB  
Article
Research on the Accumulative Damage of Flywheels Due to In-Space Charging Effects
by Dong Tian, Yanjun Feng, Hongbo Su, Xiao Zeng, Gang Liu, Yenan Liu and Jing He
Aerospace 2025, 12(2), 98; https://doi.org/10.3390/aerospace12020098 - 28 Jan 2025
Viewed by 324
Abstract
High-speed rotating flywheel bearings, designed for space applications, generate a high-resistance hydrodynamic lubrication film, which isolates the rotor, transforming it into a conductor. This phenomenon introduces a novel failure mode—flywheel bearing electrical damage caused by space charging effects. This paper first reviews the [...] Read more.
High-speed rotating flywheel bearings, designed for space applications, generate a high-resistance hydrodynamic lubrication film, which isolates the rotor, transforming it into a conductor. This phenomenon introduces a novel failure mode—flywheel bearing electrical damage caused by space charging effects. This paper first reviews the sources of common shaft voltages in flywheels and the mechanisms of electrical damage and improves the principle of deep charge causing shaft voltages in flywheel bearings, proposing that surface charge is another source of shaft voltages. The quantified analysis model of flywheel bearing electrical damage in relation to rotational speed and high-energy electron flux is derived, indicating that the damage caused by space charge–discharge to the bearing is of small magnitude and only becomes apparent after long-term accumulation, thus being easily overlooked. Based on the causal chain of electrical damage, a correlation analysis model consistent with physical principles is constructed, and the correlation between on-orbit anomalies of the flywheel and high-energy electron flux is confirmed through the use of big data. Preliminary experiments are conducted to validate all of the research results. Finally, suggestions are given for the reliable design, application, and testing of flywheels. Full article
(This article belongs to the Section Astronautics & Space Science)
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<p>Example of shaft voltage and shaft current caused by magnetic asymmetry. ((<b>a</b>) the axis is not the axis of the rotating magnetic field (i.e., magnetic asymmetry); (<b>b</b>) the axis will generate a potential difference due to cutting alternating magnetic field lines, namely the shaft voltage).</p>
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<p>Schematic diagram of two-level variable frequency drive motor circuit during the start-up phase.</p>
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<p>Common parasitic capacitance in variable frequency motors (example).</p>
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<p>Equivalent circuit of high-frequency common mode current.</p>
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<p>Shaft voltage and current caused by frequency converter power supply.</p>
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<p>High-energy electron flux and nominal E-field strength vs. driving current exceeding the standard.</p>
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<p>Distribution of nominal E-field strength vs. excessive driving current for a certain momentum wheel bearing.</p>
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<p>Summary of nominal E-field strength vs. excessive driving current for multiple momentum wheel bearings. (the red curve represents the maximum value of the driving current, and the black dashed line is the fitted result).</p>
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19 pages, 11019 KiB  
Article
Study of the Effect of Graphene Content on the Electrical and Mechanical Properties of Aluminium–Graphene Composites
by Beata Smyrak and Marek Gniełczyk
Materials 2025, 18(3), 590; https://doi.org/10.3390/ma18030590 - 28 Jan 2025
Viewed by 594
Abstract
The present paper is dedicated to the search for an alternative material based on an aluminum (Al)—few-layer graphene (FLG) composite for use in electrical applications. Due to its excellent properties, graphene has the potential for use in many applications, especially in electronics, electrical [...] Read more.
The present paper is dedicated to the search for an alternative material based on an aluminum (Al)—few-layer graphene (FLG) composite for use in electrical applications. Due to its excellent properties, graphene has the potential for use in many applications, especially in electronics, electrical engineering, aerospace, and the automotive industry. One area where the properties of graphene can be exploited is in overhead power transmission, where the main challenge at the moment is to reduce transmission losses. The utilization of conductors that exhibit superior electrical conductivity is instrumental in ensuring the mitigation of transmission losses. The utilization of graphene or other carbon allotropes is appealing due to their elevated electrical conductivity, substantial mechanical strength, and considerable heat resistance, which can enhance the properties of the composite, thereby increasing its resistance to operational conditions, particularly long-term exposure to temperature, a parameter closely related to the current carrying capacity of the OHL. This article presents the findings of research on the production of a composite based on aluminum powder and graphene, as well as the identification of its electrical and mechanical properties. The primary challenge in this research lies in the development of a method to synthesize carbon materials with aluminum using powder metallurgy, with particular attention paid to the mixing and compacting process, which is of significant importance in ensuring the appropriate distribution of carbon material in the composite. The research carried out has determined the influence of the graphene content (0.1–1 wt.%) on the electrical conductivity (max. 35.4 MS/m) and mechanical properties of Al-FLG composites (UTS = 156 MPa). Full article
(This article belongs to the Special Issue Mechanical Behavior of Composite Materials (3rd Edition))
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<p>Compilation of scanning electron microscope images of graphene used in the synthesis: (<b>a</b>) surface, (<b>b</b>) at a magnification of ×500, and (<b>c</b>) at a magnification of ×50,000.</p>
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<p>Raman spectrum that reveals the characteristics of the carbon material.</p>
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<p>The SEM images of aluminum powder used in the synthesis: (<b>a</b>) surface of aluminum powder; (<b>b</b>,<b>c</b>) SEM images of aluminum powder surface; (<b>b</b>)—magnification: 50×; (<b>c</b>)—magnification: 1000×.</p>
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<p>The SEM image of aluminum powder, with grain size measurements, magnification 2000×.</p>
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<p>Schematic representation of the manufacturing of aluminum–graphene composites.</p>
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<p>Preparing the appropriate FLG content and introducing it into aluminum powder, (<b>a</b>) FLG, (<b>b</b>) aluminum powder, (<b>c</b>) mixing of FLG and aluminum powder.</p>
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<p>Turbula mixer for mixing process, (<b>a</b>)—general view of the device, (<b>b</b>)—internal view of the device.</p>
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<p>Extrusion with rotating die: (1) press ram, (2) recipient sleeve, (3) cyclically rotating die, (4) feedstock material, and (5) finished product.</p>
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<p>SEM analysis of the distribution of carbon material in Al-FLG (0.2 wt.%), powder mixture (<b>a</b>,<b>b</b>) and EDS analysis (<b>c</b>), magnification: ×50.</p>
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<p>Localization of graphene for WDS analysis of the Al-FLG (0.2 wt.%) powder mixture (<b>left</b>) and the results of the chemical composition analysis of the powder mixture Al-FLG (0.2 wt.%) (<b>right</b>)—measurement no 1.</p>
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<p>Localization of graphene for WDS analysis of the Al-FLG (0.2 wt.%) powder mixture (<b>left</b>) and results of the chemical composition analysis of the powder mixture Al–graphene (0.2 wt.%) (<b>right</b>)—measurement no 2.</p>
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<p>Localization of points for chemical composition measurement using the EDS method (<b>on the left</b>); images of the characteristic spectrum of powder mixture Al-FLG (0.2 wt.%), magnification 2000×.</p>
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<p>Elemental distribution maps for region No. 2, Al-FLG (0.2 wt.%), magnification 2000×; (<b>a</b>) sulfur, (<b>b</b>) aluminum, (<b>c</b>) chromium, (<b>d</b>) oxygen, (<b>e</b>) copper, (<b>f</b>) carbon, (<b>g</b>) sulfur, (<b>h</b>) iron.</p>
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<p>SEM analysis of the distribution of carbon material in Al-FLG (1.0 wt.%), powder mixture (<b>a</b>,<b>b</b>) and EDS analysis (<b>c</b>), magnification: ×50.</p>
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<p>Localization of carbon material and WDS chemical composition analysis of the Al-FLG (1.0 wt.%) powder mixture (<b>on the left</b>) and the results of the chemical composition analysis of the Al-FLG (1.0 wt.%) powder mixture for measurement No. 2 (<b>on the right</b>).</p>
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<p>Localization of points for chemical composition measurements using the EDS method magnification ×2000. (<b>left</b>) and images of the characteristic spectrum, area no. 1, 2 and 3 of the Al-FLG (1.0 wt.%) mixture (<b>right</b>).</p>
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<p>Element distribution map for region no. 1, powder mixture no. 2 Al-FLG (1.0 wt.%), surface 2000×: (<b>a</b>) SL, (<b>b</b>) Al., (<b>c</b>) Cr, (<b>d</b>) O, (<b>e</b>) CP, (<b>f</b>) C, (<b>g</b>) S, and (<b>h</b>) Fe.</p>
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<p>The compacted samples of Al-FLG composites with graphene content of (<b>a</b>) 0.1%, (<b>b</b>) 0.2%, (<b>c</b>) 0.5%, and (<b>d</b>) 1 wt.%.</p>
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<p>Surface of the aluminum–graphene rods obtained in the extrusion process.</p>
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<p>The results of the electrical conductivity of the Al-graphene composites obtained in this work against the results presented in the literature [<a href="#B20-materials-18-00590" class="html-bibr">20</a>,<a href="#B30-materials-18-00590" class="html-bibr">30</a>].</p>
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<p>UTS results of Al-graphene composites obtained in this work against the results presented in the literature [<a href="#B10-materials-18-00590" class="html-bibr">10</a>,<a href="#B13-materials-18-00590" class="html-bibr">13</a>,<a href="#B17-materials-18-00590" class="html-bibr">17</a>,<a href="#B20-materials-18-00590" class="html-bibr">20</a>,<a href="#B21-materials-18-00590" class="html-bibr">21</a>,<a href="#B22-materials-18-00590" class="html-bibr">22</a>,<a href="#B31-materials-18-00590" class="html-bibr">31</a>,<a href="#B32-materials-18-00590" class="html-bibr">32</a>,<a href="#B33-materials-18-00590" class="html-bibr">33</a>,<a href="#B34-materials-18-00590" class="html-bibr">34</a>].</p>
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<p>SEM images of the pure aluminum rod fracture surface after uniaxial tensile testing; magnification: (<b>a</b>) ×100 SE, (<b>b</b>) ×200 SE, (<b>c</b>) ×1000 SE.</p>
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<p>SEM images of the Al-FLG (0.2 wt.%) composite rod fracture surface graphene content; 0.2% magnification; (<b>a</b>) ×50, SE, (<b>b</b>) ×1000 PDBSE, (<b>c</b>) ×10,000 SE.</p>
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<p>SEM images of the Al-FLG (0.5 wt.%) composite rod fracture surface; magnification: (<b>a</b>) ×50, SE, (<b>b</b>) ×1000 PDBSE, (<b>c</b>) ×10,000 SE.</p>
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<p>SEM images of the surface of the Al-FLG (1 wt.%); magnification: (<b>a</b>) ×50, SE, (<b>b</b>) ×1000 PDBSE, (<b>c</b>) ×10,000 SE.</p>
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17 pages, 3226 KiB  
Article
Investigation into the Prediction of the Service Life of the Electrical Contacting of a Wheel Hub Drive
by Markus Hempel, Niklas Umland and Matthias Busse
World Electr. Veh. J. 2025, 16(2), 68; https://doi.org/10.3390/wevj16020068 - 25 Jan 2025
Viewed by 246
Abstract
This article examines contacting by means of ultrasonic welding between a cast aluminum winding and a copper conductor of a wheel hub drive for a passenger car. The effect of thermal stress on the formation and growth of intermetallic phases (IMC) in the [...] Read more.
This article examines contacting by means of ultrasonic welding between a cast aluminum winding and a copper conductor of a wheel hub drive for a passenger car. The effect of thermal stress on the formation and growth of intermetallic phases (IMC) in the contact is analyzed. By using microscopy, the growth constant under the specific load conditions can be identified with the help of the parabolic time law and offer a possibility for predicting the service life of the corresponding contacts. As a result, it can be stated that the increase in electrical resistance of the present contact at load temperatures of 120 °C, 150 °C, and 180 °C does not reach a critical value. The growth rates of the IMC also show no critical tendencies at the usual operating temperatures (120 °C and 150 °C, e.g., at 150 °C = 4.59 × 10−7 μm2/s). The activation energy calculated using the Arrhenius plot of 155 kJ/mol (1.61 eV) can be classified as high in comparison to similar studies. In addition, it was found that future investigations of the IMC growth of corresponding electrical contacts should rather be carried out with electric current. The 180 °C sample series were carried out in the oven and with electric current; the samples in the oven did not show clear IMC, while the samples exposed to electric current already showed IMC under the microscope. Full article
17 pages, 3504 KiB  
Article
Discussion on AC Resistance and Temperature of ACSR Based on Finite Element Model Assistance
by Jianbo Yu, Changqing Wu, Hao Huang, Dexin Xie, Feixiang Qin, Jian Jiang and Gaohui He
Energies 2025, 18(3), 539; https://doi.org/10.3390/en18030539 - 24 Jan 2025
Viewed by 284
Abstract
In overhead wire transmission systems, the presence of AC resistance results in increased energy dissipation, adversely affecting the lines’ capacity to conduct current. This paper employs a finite element aluminum conductor steel-reinforced (ACSR) model, combined with electrical measurement techniques, to investigate AC resistance. [...] Read more.
In overhead wire transmission systems, the presence of AC resistance results in increased energy dissipation, adversely affecting the lines’ capacity to conduct current. This paper employs a finite element aluminum conductor steel-reinforced (ACSR) model, combined with electrical measurement techniques, to investigate AC resistance. By applying varying levels of AC current, the model is employed to determine the AC resistance which closely aligns with theoretical values estimated using the Morgan algorithm. The trends observed in the parameters are consistent, thereby validating the accuracy of the model. Following simulations and analyses regarding both AC resistance and temperature variations within the conductors—and incorporating empirical measurement results—it is demonstrated that, when environmental factors are not considered, any increase in the conductor temperature can be integrated into a revised model. This updated model is subsequently compared against test results obtained from an experimental platform; the findings confirm that the estimation errors remain within an acceptable range. Overall, this simulation model serves as a valuable reference for assessing AC losses in existing conductors, as well as contributing to reduced experimental costs while mitigating the associated risks and challenges. In summary, this simulation model serves as an essential reference for assessing AC losses in current conductors and aids in reducing experimental costs while addressing the associated risks and challenges. Full article
(This article belongs to the Special Issue Power Cables in Energy Systems)
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<p>Refinement of the experimental platform setup: (<b>a</b>) the structure of a transmission line tower and (<b>b</b>) 60 m tested ACSR.</p>
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<p>Experimental circuit wiring diagram.</p>
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<p>Three-dimensional modeling diagram: (<b>a</b>) two-layer steel core radius <math display="inline"><semantics> <mrow> <msub> <mi mathvariant="normal">R</mi> <mrow> <mi>steel</mi> </mrow> </msub> </mrow> </semantics></math> and three-layer aluminum stranded wire <math display="inline"><semantics> <mrow> <msub> <mi mathvariant="normal">R</mi> <mrow> <mi>Al</mi> </mrow> </msub> </mrow> </semantics></math> and (<b>b</b>) measured pitch length.</p>
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<p>Graph typing grid: (<b>a</b>) end face (extremely fine precision) and (<b>b</b>) side sweep (number of units 60).</p>
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<p>Finite element simulation of temperature and resistivity under five AC currents: (<b>a</b>) JL3/G1A 300/25 and (<b>b</b>) JL/G1A 400/35.</p>
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<p>Analysis chart of error bands: (<b>a</b>) JL/G1A 300/25; (<b>b</b>) JL3/G1A 300/25; and (<b>c</b>) JL/G1A 400/35.</p>
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<p>Measured temperature curves.</p>
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<p>Simulation temperature fitting curves.</p>
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<p>The steady-state temperatures rise fitting curves.</p>
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<p>Analysis chart of error bands after temperature revision: (<b>a</b>) JL/G1A 300/25; (<b>b</b>) JL3/G1A 300/25; and (<b>c</b>) JL/G1A 400/35.</p>
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<p>Analysis chart of error bands after temperature revision: (<b>a</b>) JL/G1A 300/25; (<b>b</b>) JL3/G1A 300/25; and (<b>c</b>) JL/G1A 400/35.</p>
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28 pages, 14486 KiB  
Article
Hollow Direct Air-Cooled Rotor Windings: Conjugate Heat Transfer Analysis
by Avo Reinap, Samuel Estenlund and Conny Högmark
Machines 2025, 13(2), 89; https://doi.org/10.3390/machines13020089 - 23 Jan 2025
Viewed by 326
Abstract
This article focuses on the analysis of a direct air-cooled rotor winding of a wound field synchronous machine, the innovation of which lies in the increase in the internal cooling surface, the cooling of the winding compared to the conventional inter-pole cooling, and [...] Read more.
This article focuses on the analysis of a direct air-cooled rotor winding of a wound field synchronous machine, the innovation of which lies in the increase in the internal cooling surface, the cooling of the winding compared to the conventional inter-pole cooling, and the development of a CHT evaluation model accordingly. Conjugate heat transfer (CHT) analysis is used to explore the cooling efficacy of a parallel-cooled hollow-conductor winding of a salient-pole rotor and to identify a cooling performance map. The use of high current densities of 15–20 Arms/mm2 in directly cooled windings requires high cooling intensity, which in the case of air cooling results not only in flow velocities above 15 m/s to ensure permissible operating temperatures, but also the need for coolant distribution and heat transfer studies. The experiments and calculations are based on a non-rotating machine and a wind tunnel using the same rotor coil(s). CHT-based thermal calculations provide not only reliable results compared to experimental work and lumped parameter thermal circuits with adjusted aggregate parameters, but also insight related to pressure and cooling flow distribution, thermal loads, and cooling integration issues that are necessary for the development of high power density and reliable electrical machines. The results of the air-cooling integration show that the desired high current density is achievable at the expense of high cooling intensity, where the air velocity ranges from 15 to 30 m/s and 30 to 55 m/s, distinguishing the air velocity of the hollow conductor and bypass channel, compared to the same coil in an electric machine and a wind tunnel at the similar thermal load and limit. Since the hot spot location depends on cooling integration and cooling intensity, modeling and estimating the cooling flow is essential in the development of wound-field synchronous machines. Full article
(This article belongs to the Section Electrical Machines and Drives)
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<p>Overview of the test machine [<a href="#B18-machines-13-00089" class="html-bibr">18</a>]: (<b>a</b>) Longitudinal cross-section of the electric machine. From left to right: slip-ring system, cooling system inlet interface (blue highlighted), test machine, and mounting flange. (<b>b</b>) The outlet of the test machine makes visible both the rotor windings and the placed thermocouples.</p>
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<p>IR image at stand-still rotor thermal test 120 A, 1.6 m<sup>3</sup>/min, and inlet temperature 40 °C.</p>
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<p>Thermal test values of a stationary machine with increased rotor current (J<sub>max</sub> = 16.2 A/mm<sup>2</sup>).</p>
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<p>Presentation of 3 + 1 cross sections. A group of 3: symmetrical half-cross section of the rotor pole and cooling/bypass channels (<b>left</b>), markings used for velocity, pressure, and temperature distribution along the rotor (<b>middle</b>), and thermal resistances and unforced cooling interfaces (<b>right</b>). Rightmost figure: designation of the cooling channels in the wind tunnel.</p>
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<p>2D FE heat transfer evaluation of a wind tunnel experiment (<b>left</b>), base geometry for 3D FE heat transfer evaluation (<b>middle</b>), and a section of CAD drawing for machine prototype (<b>right</b>).</p>
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<p>Half of the rotor pole. The color bars show the velocity, pressure, and temperature distribution along the rotor winding at Q<sub>in</sub> = 1.36 m<sup>3</sup>/min (170 L/min per coil) and P<sub>rcu</sub> = 1.39 kW: (<b>a</b>) inlet temperature ϑ<sub>in</sub> = 20 °C (J = 20.2 A/mm<sup>2</sup>); (<b>b</b>) inlet temperature ϑ<sub>in</sub> = 40 °C (J = 19.6 A/mm<sup>2</sup>).</p>
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<p>Half of the rotor pole. The color bars show the velocity, pressure and temperature distribution along the rotor winding Q<sub>in</sub> = 0.68 m<sup>3</sup>/min (85 L/min per coil) and P<sub>rcu</sub> = 1.39 kW: (<b>a</b>) inlet temperature ϑ<sub>in</sub> = 20 °C (J = 18.5 A/mm<sup>2</sup>); (<b>b</b>) inlet temperature ϑ<sub>in</sub> = 40 °C (J = 17.9 A/mm<sup>2</sup>).</p>
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<p>Average outlet flow rate (Q<sub>out</sub>), inlet pressure (p<sub>in</sub>), maximum flow velocity in the cooling ducts (u<sub>max</sub>), average outlet temperature (ϑ<sub>out</sub>), minimum (ϑ<sub>wmin</sub>) and maximum coil temperature (ϑ<sub>wmax</sub>) as a function of heating power (P<sub>rcu</sub>), and inlet flow rate (Q<sub>in</sub>). The inlet temperatures are 20 °C and 40 °C, shown by thin and thick contour lines, respectively.</p>
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<p>Location of temperature sensors (<b>left</b>) and IR image at the 18th minute of the test (<b>right</b>).</p>
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<p>Presentation of the selected parameters of the heat test as measured experimentally on the wind tunnel (J<sub>max</sub> = 20.2 A/mm<sup>2</sup>).</p>
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<p>Distribution of pressure, flow rate, flow temperature, and conductor temperature along the cooling channels in wind-tunnel 3D FE CHT model. The data are labeled as shown in <a href="#machines-13-00089-f004" class="html-fig">Figure 4</a>.</p>
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<p>Half of the rotor coil in wind tunnel. The color bars show the velocity, pressure and temperature distribution at Q<sub>in</sub> = 1.36 m<sup>3</sup>/min (170 L/min per coil) and P<sub>rcu</sub> = 1.39 kW: (<b>a</b>) Inlet temperature ϑ<sub>in</sub> = 20 °C (J = 20.4 A/mm<sup>2</sup>); (<b>b</b>) Inlet temperature ϑ<sub>in</sub> = 40 °C (J = 19.8 A/mm<sup>2</sup>).</p>
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<p>Half of the rotor coil in wind tunnel. The color bars show the velocity, pressure, and temperature distribution at Q<sub>in</sub> = 0.68 m<sup>3</sup>/min (85 L/min per coil) and P<sub>rcu</sub> = 1.39 kW. (<b>a</b>) Inlet temperature ϑ<sub>in</sub> = 20 °C (J = 18.5 A/mm<sup>2</sup>); (<b>b</b>) inlet temperature ϑ<sub>in</sub> = 40 °C (J = 18.1A/mm<sup>2</sup>).</p>
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<p>Average outlet flow rate (Q<sub>out</sub>), inlet pressure (p<sub>in</sub>), maximum flow velocity in the cooling ducts (u<sub>max</sub>), average outlet temperature (ϑ<sub>out</sub>), and minimum (ϑ<sub>wmin</sub>) and maximum coil temperature (ϑ<sub>wmax</sub>) as a function of heating power (P<sub>rcu</sub>) and inlet flow rate (Q<sub>in</sub>). The contour lines show ϑ<sub>in</sub> = 20 °C (thin), ϑ<sub>in</sub> = 40 °C (thick), and the effect of inter-turn thermal resistance (dotted).</p>
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<p>Distribution of pressure, flow rate, flow temperature, and conductor temperature along the cooling channels (<a href="#machines-13-00089-f004" class="html-fig">Figure 4</a>) in the rotor-pole model without thermal resistance between the turns.</p>
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<p>Distribution of pressure, flow rate, flow temperature, and conductor temperature along the cooling channels (<a href="#machines-13-00089-f004" class="html-fig">Figure 4</a>) in the wind-tunnel model without thermal resistance between the turns.</p>
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<p>Distribution of pressure, flow rate, flow temperature, and conductor temperature along the cooling channels (<a href="#machines-13-00089-f004" class="html-fig">Figure 4</a>) in the wind-tunnel model with thermal resistance between the turns. The temperatures of the winding conductors are divided into two groups: (1) linear—hollow conductor and bypass cooling and (2) parabolic—hollow conductor cooling only.</p>
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<p>Distribution heat transfer coefficient and heat flux along the cooling channels (<a href="#machines-13-00089-f004" class="html-fig">Figure 4</a>) in the rotor pole model without thermal resistance between the turns.</p>
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<p>Distribution heat transfer coefficient and heat flux along the cooling channels (<a href="#machines-13-00089-f004" class="html-fig">Figure 4</a>) in the wind-tunnel model with thermal resistance between the turns.</p>
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<p>Distribution of temperature measurement points along the test coil in the wind tunnel and their relationship to LPTN model-based estimates of the copper and air temperatures of the hollow conductors and bypass channels coils.</p>
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<p>Temperature distribution along the rotor coil, rotor teeth, and cooling channels at 1600 L/min and 140 A and 2400 L/min at 180 A and 197 A (J = 22.5A/mm<sup>2</sup>), respectively.</p>
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<p>Lumped circuit model used in Motor-CAD. The thermal power across the rotor coil top and coil side thermal resistances are shown per axial slice of 3 slices.</p>
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<p>Temperature distribution: (<b>a</b>) axial 2D FEA; (<b>b</b>) rotor pole radial FEA.</p>
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<p>Heat transfer at the inner-wall surfaces of a hollow conductor for a wind tunnel model of a rotor coil at Q<sub>in</sub> = 1.36 m<sup>3</sup>/min, P<sub>rcu</sub> = 1.39kW (J = 19.8 A/mm<sup>2</sup>), ϑ<sub>in</sub> = 40 °C and inter-coil thermal resistance of 10 W/(m<sup>2</sup> K): (<b>a</b>) section and surface temperature [°C] and temperature gradient and flow vectors; (<b>b</b>) normal component of heat flux [W/m<sup>2</sup>] on wall surfaces, transversal component of total heat flux on winding sections, including vectors.</p>
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<p>Heat transfer at the inner-wall surfaces of a hollow conductor for a rotor pole model of a rotor coil at Q<sub>in</sub> = 1.36 m<sup>3</sup>/min, P<sub>rcu</sub> = 1.39kW (J = 19.6 A/mm<sup>2</sup>), ϑ<sub>in</sub> = 40 °C and inter-coil thermal resistance of 10 W/(m<sup>2</sup> K): (<b>a</b>) section and surface temperature [°C] and temperature gradient and flow vectors; (<b>b</b>) normal component of heat flux [W/m<sup>2</sup>] on wall surfaces, transversal component of total heat flux on winding sections, including vectors.</p>
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<p>A brief overview of the calculations and measurements of the maximum temperature of the rotor coil (calculation points shown with *). On the left is the contour map of the rotor coil model, and on the right is the wind tunnel (calculation points shown with x).</p>
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<p>Flux density distribution at 120, 140, 160, 180, and 200 A of excitation current.</p>
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24 pages, 12257 KiB  
Article
Fast Simulation of Electromagnetic Scattering for Radar-Absorbing Material-Coated 3D Electrically Large Targets
by Hongzu Li, Chunlei Dong, Lixin Guo, Xiao Meng and Dan Wang
Remote Sens. 2025, 17(3), 390; https://doi.org/10.3390/rs17030390 - 23 Jan 2025
Viewed by 465
Abstract
In this paper, a modified Shooting and Bouncing Ray (SBR) method based on high-order impedance boundary conditions (HOIBCs) is proposed to analyze the electromagnetic (EM) scattering from electrically large three-dimensional (3D) conducting targets coated with radar-absorbing material (RAM). In addition, the edge diffraction [...] Read more.
In this paper, a modified Shooting and Bouncing Ray (SBR) method based on high-order impedance boundary conditions (HOIBCs) is proposed to analyze the electromagnetic (EM) scattering from electrically large three-dimensional (3D) conducting targets coated with radar-absorbing material (RAM). In addition, the edge diffraction field of coated targets is included in the calculation to improve the accuracy of the calculation. Firstly, the SBR method based on the bidirectional tracing technique is presented. It is concluded that the calculation of the scattered field of the coated targets requires the determination of the reflection coefficients on the coated surface. The reflection coefficients of the coated targets are then derived using HOIBC theory. Finally, the equivalent edge current (EEC) of the impedance wedge is derived by integrating the UTD solutions for the impedance wedge diffraction with the impedance boundary conditions. The simulation results show that the proposed method improves computational efficiency compared to MLFMA while maintaining accuracy. Furthermore, the RCS characteristics of targets coated with different RAMs, different coating thicknesses and with different angles of incidence were compared, as well as the RCS results of coated targets with those of conventional perfect electrical conductor (PEC) targets. Full article
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Figure 1

Figure 1
<p>(<b>a</b>) Multiple reflections of ray tubes; (<b>b</b>) schematic diagram of the bidirectional tracing technique.</p>
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<p>(<b>a</b>) Polarization diagram of TE wave illumination; (<b>b</b>) polarization diagram of TM wave illumination.</p>
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<p>Metallic conductor plate coated with a dielectric layer.</p>
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<p>Impedance wedge geometric model. (<b>a</b>) Overall schematic diagram of the wedge structure. (<b>b</b>) Schematic diagram along the wedge edge. (<b>c</b>) Cross-sectional schematic diagram of the wedge.</p>
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<p>Illumination region decomposition.</p>
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<p>Flowchart of the proposed method.</p>
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<p>Monostatic RCS of the dihedral reflector. (<b>a</b>) VV polarization. (<b>b</b>) HH polarization.</p>
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<p>Monostatic RCS of the cube. (<b>a</b>) HH polarization. (<b>b</b>) VV polarization.</p>
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<p>Monostatic RCS of the coated composite cube structure. (<b>a</b>) Monostatic RCS for RMA1 with HH polarization. (<b>b</b>) Monostatic RCS for RMA2 with HH polarization. (<b>c</b>) Monostatic RCS for RMA1 with VV polarization. (<b>d</b>) Monostatic RCS for RMA2 with VV polarization.</p>
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<p>Monostatic RCS of the coated composite cube structure. (<b>a</b>) Monostatic RCS for RMA1 with HH polarization. (<b>b</b>) Monostatic RCS for RMA2 with HH polarization. (<b>c</b>) Monostatic RCS for RMA1 with VV polarization. (<b>d</b>) Monostatic RCS for RMA2 with VV polarization.</p>
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<p>Model of aircraft.</p>
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<p>Monostatic RCS on the <span class="html-italic">xoz</span>-plane of aircraft model at 10 GHz. (<b>a</b>) VV polarization. (<b>b</b>) HH polarization.</p>
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<p>Monostatic RCS on the <span class="html-italic">xoy</span>-plane of aircraft model at 10 GHz. (<b>a</b>) VV polarization. (<b>b</b>) HH polarization.</p>
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<p>Monostatic RCS with different thicknesses on the coated composite cube structure. (<b>a</b>) Thickness is <math display="inline"><semantics> <mrow> <mn>0.015</mn> <mi>λ</mi> </mrow> </semantics></math> with HH polarization. (<b>b</b>) Thickness is <math display="inline"><semantics> <mrow> <mn>0.075</mn> <mi>λ</mi> </mrow> </semantics></math> with HH polarization. (<b>c</b>) Thickness is <math display="inline"><semantics> <mrow> <mn>0.015</mn> <mi>λ</mi> </mrow> </semantics></math> with VV polarization. (<b>d</b>) Thickness is <math display="inline"><semantics> <mrow> <mn>0.075</mn> <mi>λ</mi> </mrow> </semantics></math> with VV polarization.</p>
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<p>Monostatic RCS with different thicknesses on the coated composite cube structure. (<b>a</b>) Thickness is <math display="inline"><semantics> <mrow> <mn>0.015</mn> <mi>λ</mi> </mrow> </semantics></math> with HH polarization. (<b>b</b>) Thickness is <math display="inline"><semantics> <mrow> <mn>0.075</mn> <mi>λ</mi> </mrow> </semantics></math> with HH polarization. (<b>c</b>) Thickness is <math display="inline"><semantics> <mrow> <mn>0.015</mn> <mi>λ</mi> </mrow> </semantics></math> with VV polarization. (<b>d</b>) Thickness is <math display="inline"><semantics> <mrow> <mn>0.075</mn> <mi>λ</mi> </mrow> </semantics></math> with VV polarization.</p>
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<p>Monostatic RCS on the <span class="html-italic">xoz</span>-plane with different thicknesses on the aircraft. (<b>a</b>) VV polarization. (<b>b</b>) HH polarization.</p>
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<p>Monostatic RCS on the <span class="html-italic">xoy</span>-plane with different thicknesses on the aircraft. (<b>a</b>) VV polarization. (<b>b</b>) HH polarization.</p>
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<p>Bistatic RCS of two balls at <math display="inline"><semantics> <mrow> <msub> <mi>θ</mi> <mi>i</mi> </msub> <mo>=</mo> <msup> <mn>0</mn> <mo>∘</mo> </msup> <mo>,</mo> <mi>φ</mi> <mo>=</mo> <msup> <mn>0</mn> <mo>∘</mo> </msup> </mrow> </semantics></math>. (<b>a</b>) VV polarization. (<b>b</b>) HH polarization.</p>
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<p>Bistatic RCS of the aircraft at <math display="inline"><semantics> <mrow> <msub> <mi>θ</mi> <mi>i</mi> </msub> <mo>=</mo> <msup> <mrow> <mn>30</mn> </mrow> <mo>∘</mo> </msup> <mo>,</mo> <mi>φ</mi> <mo>=</mo> <msup> <mn>0</mn> <mo>∘</mo> </msup> </mrow> </semantics></math>. (<b>a</b>) VV polarization. (<b>b</b>) HH polarization.</p>
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<p>Monostatic RCS of the aircraft at <math display="inline"><semantics> <mrow> <msub> <mi>θ</mi> <mi>i</mi> </msub> <mo>=</mo> <mo>−</mo> <msup> <mrow> <mn>30</mn> </mrow> <mo>∘</mo> </msup> <mo>,</mo> <mi>φ</mi> <mo>=</mo> <msup> <mn>0</mn> <mo>∘</mo> </msup> </mrow> </semantics></math>. (<b>a</b>) VV polarization. (<b>b</b>) HH polarization.</p>
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<p>Monostatic RCS of the aircraft at <math display="inline"><semantics> <mrow> <msub> <mi>θ</mi> <mi>i</mi> </msub> <mo>=</mo> <mo>−</mo> <msup> <mrow> <mn>120</mn> </mrow> <mo>∘</mo> </msup> <mo>,</mo> <mi>φ</mi> <mo>=</mo> <msup> <mn>0</mn> <mo>∘</mo> </msup> </mrow> </semantics></math>. (<b>a</b>) VV polarization. (<b>b</b>) HH polarization.</p>
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<p>Monostatic RCS of the aircraft at <math display="inline"><semantics> <mrow> <msub> <mi>θ</mi> <mi>i</mi> </msub> <mo>=</mo> <msup> <mrow> <mn>120</mn> </mrow> <mo>∘</mo> </msup> <mo>,</mo> <mi>φ</mi> <mo>=</mo> <msup> <mn>0</mn> <mo>∘</mo> </msup> </mrow> </semantics></math>. (<b>a</b>) VV polarization. (<b>b</b>) HH polarization.</p>
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<p>Model of ship.</p>
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<p>Monostatic RCS on the <span class="html-italic">xoy</span>-plane of the ship model at 6 GHz. (<b>a</b>) VV polarization. (<b>b</b>) HH polarization.</p>
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<p>Monostatic RCS on the <span class="html-italic">xoy</span>-plane of the ship model at 8 GHz. (<b>a</b>) VV polarization. (<b>b</b>) HH polarization.</p>
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<p>Monostatic RCS on the <span class="html-italic">xoy</span>-plane of the ship model at 10 GHz. (<b>a</b>) VV polarization. (<b>b</b>) HH polarization.</p>
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40 pages, 17765 KiB  
Article
Aerodynamic and Vibration Characteristics of Iced Power Transmission Conductors in a Nonuniform Wind Field Based on Unsteady Theory
by Guifeng Zhao, Qingyang Li, Xiuyao Li and Meng Zhang
Energies 2025, 18(3), 459; https://doi.org/10.3390/en18030459 - 21 Jan 2025
Viewed by 366
Abstract
To study the aerodynamic and vibration characteristics of iced conductors under the influence of wind fluctuations, a harmonic superposition method is used to simulate nonuniform wind speeds. A user-defined function is written on the basis of the secondary development function of the Fluent [...] Read more.
To study the aerodynamic and vibration characteristics of iced conductors under the influence of wind fluctuations, a harmonic superposition method is used to simulate nonuniform wind speeds. A user-defined function is written on the basis of the secondary development function of the Fluent 2021 R1 software to determine the displacement and velocity of the conductor at each time step, and a two-way fluid–structure interaction (FSI) numerical simulation of an iced conductor under a nonuniform wind field is performed via an overset mesh method. In the analysis, the aerodynamic coefficients and galloping characteristics of iced conductors under different degrees of freedom (DOFs) are investigated by considering different combinations of quasi-steady theory, unsteady theory, a uniform wind field, and a nonuniform wind field. The results show that in a nonuniform wind field, the mean, standard deviation (SD), and peak values of the drag and torsion coefficients of the conductors calculated via unsteady theory are significantly larger than those calculated via quasi-steady theory, indicating that the obtained aerodynamic coefficients of the latter (the mean values are typically used) conceal the characteristics of the iced conductors in an actual wind environment and ignore the adverse effects of the variability. Full article
(This article belongs to the Special Issue Advances in Fluid Dynamics and Wind Power Systems: 2nd Edition)
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Figure 1

Figure 1
<p>Dynamic model of an iced conductor under different DOFs. (<b>a</b>) one-way FSI model (<b>b</b>) 1-DOF model (<b>c</b>) 3-DOFs model (<b>d</b>) schematic diagram of aerodynamic force and wind attack angle on iced conductor.</p>
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<p>Calculation flow chart under unsteady theory.</p>
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<p>Schematic diagram of the iced conductor cross-section and the calculation domain of the flow field. (<b>a</b>) bare conductor (<b>b</b>) conductor with the crescent-shaped iced section (<b>c</b>) computational domain.</p>
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<p>The specific forms of the foreground and background grids. (<b>a</b>) background grid (<b>b</b>) foreground grid (<b>c</b>) data-exchange layer.</p>
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<p>Vortex shedding frequencies of conductors with different ice thicknesses. (<b>a</b>) 1.1D ice thickness (<b>b</b>) 1.4D ice thickness.</p>
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<p>Numerical simulation results of fluctuating wind speed. (<b>a</b>) simulated wind speed spectrum and target power spectrum (<b>b</b>) time–history curve of wind speed.</p>
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<p>Flowchart of nonuniform wind field realization.</p>
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<p>The organizational structure and analysis methodology are presented in <a href="#sec4-energies-18-00459" class="html-sec">Section 4</a>.</p>
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<p>Aerodynamic coefficients of iced conductors under two wind fields when quasi-steady theory is used. (<b>a</b>) mean drag coefficient (<b>b</b>) mean lift coefficient (<b>c</b>) mean moment coefficient (<b>d</b>) root mean square of drag coefficient (<b>e</b>) root mean square of lift coefficient (<b>f</b>) root mean square of moment coefficient.</p>
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<p>Den Hartog and Nigol coefficients of iced conductors under two wind fields when quasi-steady theory is used. (<b>a</b>) Den Hartog coefficient (<b>b</b>) Nigol coefficient.</p>
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<p>Time–history curves of the aerodynamic coefficients of the iced conductor with a wind attack angle of 60°. (<b>a</b>) drag coefficient (<b>b</b>) lift coefficient (<b>c</b>) moment coefficient.</p>
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<p>Aerodynamic coefficients of iced conductors calculated via different theories in a uniform wind field. (<b>a</b>) mean drag coefficient (<b>b</b>) mean lift coefficient (<b>c</b>) mean moment coefficient.</p>
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<p>Den Hartog and Nigol coefficients of iced conductors calculated via different theories in a uniform wind field. (<b>a</b>) Den Hartog coefficient (<b>b</b>) Nigol coefficient.</p>
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<p>Time–history curves of the aerodynamic coefficients of the iced conductor in a uniform wind field when the wind attack angle is 30°. (<b>a</b>) drag coefficient (<b>b</b>) lift coefficient (<b>c</b>) moment coefficient.</p>
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<p>Aerodynamic coefficients of iced conductors in a nonuniform wind field calculated via different theories. (<b>a</b>) drag coefficient (<b>b</b>) lift coefficient (<b>c</b>) moment coefficient.</p>
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<p>Time–history curves of the aerodynamic coefficients of the iced conductors in a nonuniform wind field when the wind attack angle is 30°. (<b>a</b>) drag coefficient (<b>b</b>) lift coefficient (<b>c</b>) moment coefficient.</p>
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<p>Den Hartog and Nigol coefficients of iced conductors calculated via different theories in a nonuniform wind field. (<b>a</b>) Den Hartog coefficient (<b>b</b>) Nigol coefficient.</p>
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<p>Comparison of the aerodynamic coefficients of the conductor under the two wind fields when only the vibration of the conductor in the crosswind direction is considered and unsteady theory is used. (<b>a</b>) drag coefficient (<b>b</b>) lift coefficient (<b>c</b>) moment coefficient (<b>d</b>) root mean square of drag coefficient (<b>e</b>) root mean square of lift coefficient (<b>f</b>) root mean square of moment coefficient.</p>
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<p>The Den Hartog and Nigol coefficients of the conductor are considered when unsteady theory is used, and only the vibration of the conductor in the crosswind direction is considered. (<b>a</b>) Den Hartog coefficient (<b>b</b>) Nigol coefficient.</p>
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<p>Time–history curves of the aerodynamic coefficients of the iced conductor in two wind fields calculated via unsteady theory when the wind attack angle is 30°. (<b>a</b>) drag coefficient (<b>b</b>) lift coefficient (<b>c</b>) moment coefficient.</p>
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<p>Displacement amplitudes of the conductor in the crosswind direction in the two wind fields.</p>
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<p>Time–history curves of the displacement of the conductor in the crosswind direction in the two wind fields.</p>
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<p>Aerodynamic coefficients of the iced conductors in the two wind fields when the 3-DOF-coupled vibration of the conductors is considered. (<b>a</b>) mean drag coefficient (<b>b</b>) mean lift coefficient (<b>c</b>) mean moment coefficient (<b>d</b>) root mean square of drag coefficient (<b>e</b>) root mean square of lift coefficient (<b>f</b>) root mean square of moment coefficient.</p>
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<p>Den Hartog and Nigol coefficients of iced conductors in two wind fields when the 3-DOF-coupled vibration of the conductors is considered. (<b>a</b>) Den Hartog coefficient (<b>b</b>) Nigol coefficient.</p>
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<p>Time–history curves of the aerodynamic coefficients of the iced conductor under different wind attack angles when the 3-DOF-coupled vibration of the conductors is considered. (<b>a</b>) 0° wind attack angle (<b>b</b>) 30° wind attack angle (<b>c</b>) 120° wind attack angle (<b>d</b>) 150° wind attack angle (<b>e</b>) 180° wind attack angle.</p>
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<p>Time–history curves of the aerodynamic coefficients of the iced conductor under different wind attack angles when the 3-DOF-coupled vibration of the conductors is considered. (<b>a</b>) 0° wind attack angle (<b>b</b>) 30° wind attack angle (<b>c</b>) 120° wind attack angle (<b>d</b>) 150° wind attack angle (<b>e</b>) 180° wind attack angle.</p>
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<p>Spectrum of the aerodynamic coefficients of the iced conductor under different wind attack angles. (<b>a</b>) 0° wind attack angle (<b>b</b>) 30° wind attack angle (<b>c</b>) 120° wind attack angle (<b>d</b>) 150° wind attack angle (<b>e</b>) 180° wind attack angle.</p>
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<p>Spectrum of the aerodynamic coefficients of the iced conductor under different wind attack angles. (<b>a</b>) 0° wind attack angle (<b>b</b>) 30° wind attack angle (<b>c</b>) 120° wind attack angle (<b>d</b>) 150° wind attack angle (<b>e</b>) 180° wind attack angle.</p>
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<p>Strouhal number comparison chart.</p>
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<p>Mean displacements of the iced conductors in two wind fields. (<b>a</b>) mean displacement in the crosswind direction (<b>b</b>) mean displacement in the along-wind direction (<b>c</b>) mean torsion angle.</p>
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<p>Time–history curves of conductor displacement in different wind fields when the wind attack angle is 30°. (<b>a</b>) displacement in the along-wind direction (<b>b</b>) displacement in the crosswind direction (<b>c</b>) torsion angle.</p>
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<p>Time–history curves of conductor displacement in different wind fields when the wind attack angle is 120°. (<b>a</b>) displacement in the along-wind direction (<b>b</b>) displacement in the crosswind direction (<b>c</b>) torsion angle.</p>
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<p>Displacement spectra of the iced conductor under different attack angles. (<b>a</b>) 30° wind attack angle (<b>b</b>) 120° wind attack angle.</p>
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<p>Movement trajectories of iced conductors in a nonuniform wind field calculated via unsteady theory.</p>
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<p>Velocity contours of the iced conductor in the nonuniform wind field calculated via unsteady theory. (<b>a</b>) 0° wind attack angle (<b>b</b>) 30° wind attack angle (<b>c</b>) 150° wind attack angle (<b>d</b>) 180° wind attack angle.</p>
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<p>Pressure contours of the iced conductor in the nonuniform wind field calculated via unsteady theory. (<b>a</b>) 0° wind attack angle (<b>b</b>) 30° wind attack angle (<b>c</b>) 150° wind attack angle (<b>d</b>) 180° wind attack angle.</p>
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<p>Pressure contours of the iced conductor in the nonuniform wind field calculated via unsteady theory. (<b>a</b>) 0° wind attack angle (<b>b</b>) 30° wind attack angle (<b>c</b>) 150° wind attack angle (<b>d</b>) 180° wind attack angle.</p>
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<p>Vorticity diagram of the iced conductor in the nonuniform wind field calculated via unsteady theory. (<b>a</b>) 0° wind attack angle (<b>b</b>) 30° wind attack angle (<b>c</b>) 150° wind attack angle (<b>d</b>) 180° wind attack angle.</p>
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15 pages, 4793 KiB  
Article
Dynamic Simulation of Underground Cable Laying for Digital Three-Dimensional Transmission Lines
by Chunhua Fang, Wenqi Lu, Jialiang Liu, Xiuyou Yang and Jin Zhang
Appl. Sci. 2025, 15(2), 979; https://doi.org/10.3390/app15020979 - 20 Jan 2025
Viewed by 550
Abstract
In light of the issues associated with the laying process of transmission line cables, including concealed security risks and contact collisions between pulleys and cables, which primarily stem from reliance on drawings, this paper introduces a simulation methodology for the cable laying construction [...] Read more.
In light of the issues associated with the laying process of transmission line cables, including concealed security risks and contact collisions between pulleys and cables, which primarily stem from reliance on drawings, this paper introduces a simulation methodology for the cable laying construction process utilizing Building Information Modeling (BIM) technology. Initially, two-dimensional DWG graphic data are employed to develop a model of the target equipment and construction environment using BIM software (Solid works 2020). Subsequently, the cable is accurately modeled by applying ADAMS virtual prototype technology, the bushing force connection method, and the macro command language. This allows for the construction of a three-dimensional real cable laying system for transmission lines, enabling the simulation of the dynamic cable laying process in the field. Subsequently, an error analysis is conducted to compare the axial tension and laying speed of the cable with theoretical calculation values. The study then proceeds to analyze tension fluctuations during the cable laying process and assess the load-bearing capacity of the pulleys, thus facilitating effective control of the construction process and enhancing safety measures. The findings indicate that the proposed method can accurately and efficiently simulate the on-site cable laying construction process, with numerical errors maintained below 5%, thereby validating the integrity of the model. Furthermore, the traction overload safety protection amplification coefficient is determined to be α = 1.5. It is highlighted that the bearing capacity of the block must exceed 60% of the load carried by the conductor at constant speed. This research provides a theoretical foundation for addressing safety hazards in cable laying engineering and holds certain engineering value. Full article
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<p>Target device model. (<b>a</b>) Cable tray model. (<b>b</b>) Pulley model. (<b>c</b>) Cable conveyor model.</p>
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<p>Cable tunnel environment model.</p>
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<p>Cylindrical force model of cable and traction rope.</p>
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<p>Drawing of the cable tunnel environment.</p>
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<p>Cable laying system model.</p>
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<p>Simplified diagram of cable laying path.</p>
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<p>Dynamic speed change of cable laying. (<b>a</b>) Tow rope axial velocity. (<b>b</b>) Axial velocity of cable micro-segments.</p>
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<p>Changes in dynamic axial tension of cable laying. (<b>a</b>) Axial traction of micro-segments of haul rope. (<b>b</b>) Axial tension in micro-segments of diagonal upstream cables. (<b>c</b>) Axial tension in micro-segments of horizontal section cables. (<b>d</b>) Axial tension in micro-segments of diagonal downstream cables.</p>
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<p>Errors between calculated and simulated tension values at each position.</p>
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<p>Axial tension of cable microsegment.</p>
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<p>Axial tension at the moment when the cable micro-section crosses the pulley.</p>
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<p>Ratio of maximum axial tension to theoretical calculation value.</p>
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<p>Data diagram of load borne by the pulley. (<b>a</b>) Load on pulley 1. (<b>b</b>) Load on pulley 2. (<b>c</b>) Load on pulley 3. (<b>d</b>) Load on pulley 4.</p>
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11 pages, 1773 KiB  
Communication
Current Characteristics of REBCO Tapes in 6-Slot TSTC-CICC Under Bending Conditions
by Li Li, Mingzhen Yang, Songzhen Yuan, Shaotao Dai and Tao Ma
Materials 2025, 18(2), 350; https://doi.org/10.3390/ma18020350 - 14 Jan 2025
Viewed by 386
Abstract
Embedding stacked HTS tapes into twisted slots is one design approach for constructing fusion conductors. This paper adopts a Cable-in-Conduit Conductor (CICC) structure, utilizing commercially REBCO coated conductors. The cable framework is made of copper and features six helically twisted slots filled with [...] Read more.
Embedding stacked HTS tapes into twisted slots is one design approach for constructing fusion conductors. This paper adopts a Cable-in-Conduit Conductor (CICC) structure, utilizing commercially REBCO coated conductors. The cable framework is made of copper and features six helically twisted slots filled with 2G HTS tapes. Two 1 m long samples with twist pitches of 200 mm and 300 mm, respectively, were fabricated. In one slot, copper and superconducting tapes were alternated, while the remaining grooves were filled with copper tapes. The 90 µm thick copper-plated bare tapes provided by Shanghai Superconductor were used for testing. By measuring the critical current of tapes positioned at different locations within the grooves at 77 K, the characteristics of each tape in the stacked arrangement were individually characterized. The study obtained the current degradation patterns of tapes located at different positions within the grooves under various bending radii. This paper will present and discuss the preliminary results of the bending measurements conducted at 77 K under a self-field. Full article
(This article belongs to the Special Issue Advances in Superconducting Materials for Electric Power Applications)
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<p>The figure shows the cross-section of a six-slot TSTC-CICC, where HTS tapes are located in three symmetrically positioned slots, and copper tapes are used to evenly space the tapes. The remaining slots are filled with copper tapes of the same thickness, and the outermost part is fixed by the copper sheath.</p>
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<p>The figure shows a schematic diagram for critical current testing. In the diagram, 1-1′ indicates the position where the current is introduced into the superconducting tape, 2-2′ represents the position where the voltage on the tape is measured, and the rectangular frame is the conductor after bending.</p>
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<p>A three-point bending process was conducted using a pipe bender, gradually reducing the bending radius of the conductor, and the bending radius was determined by measuring the arc length of the bent conductor. After determining the bending radius, the I–V curves of the tapes at different positions were measured.</p>
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<p>Model schematic of the REBCO layer strain during bending in a six-slot TSTC: (<b>a</b>) cross-sectional view of the tape embedded in the groove, using the bottom tape as an example; (<b>b</b>) enlarged view of a segment of the tape after twisting, embedded in the helical groove framework.</p>
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<p>The I–V curves of the tapes at different positions in the groove when the tapes are embedded in the skeleton but in an unbended state.</p>
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<p>I–V curves of tapes within the slot of the 200 mm twist pitch sample under different bending radii.</p>
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<p>I–V curves of tapes within the slot of the 300 mm twist pitch sample under different bending radii.</p>
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<p>The figure shows the analytical calculation results of the axial strain of the outermost tape in the groove under different bending radii.</p>
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<p>The figure shows the axial strain calculation results of the tape at different positions in the groove for a sample with a twist pitch of 300 mm at a bending radius of 90 cm.</p>
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14 pages, 10429 KiB  
Article
Studies of Thermal Conductivity of Graphite Foil-Based Composite Materials
by Vladimir A. Shulyak, Nikolai S. Morozov, Roman A. Minushkin, Viktor Yu. Gubin, Dmitriy V. Vakhrushin, Alexandra V. Gracheva, Ildar Kh. Nigmatullin, Sergei N. Chebotarev and Viktor V. Avdeev
Materials 2025, 18(2), 233; https://doi.org/10.3390/ma18020233 - 8 Jan 2025
Viewed by 483
Abstract
We have proposed and developed a method for measuring the thermal conductivity of highly efficient thermal conductors. The measurement method was tested on pure metals with high thermal conductivity coefficients: aluminum (99.999 wt.% Al) and copper (99.990 wt.% Cu). It was demonstrated that [...] Read more.
We have proposed and developed a method for measuring the thermal conductivity of highly efficient thermal conductors. The measurement method was tested on pure metals with high thermal conductivity coefficients: aluminum (99.999 wt.% Al) and copper (99.990 wt.% Cu). It was demonstrated that their thermal conductivities at a temperature of T = 22 ± 1 °C were <λAl> = 243 ± 3 W/m·K and <λCu> = 405 ± 4 W/m·K, which was in good agreement with values reported in the literature. Artificial graphite (ρG1 = 1.8 g/cm3) and natural graphite (ρG2 = 1.7 g/cm3) were used as reference carbon materials; the measured thermal conductivities were <λG1> = 87 ± 1 W/m·K and <λG2> = 145 ± 3 W/m·K, respectively. It is well established that measuring the thermal conductivity coefficient of thin flexible graphite foils is a complex metrological task. We have proposed to manufacture a solid rectangular sample formed by alternating layers of thin graphite foils connected by layers of ultra-thin polyethylene films. Computer modelling showed that, for equal thermal conductivities of solid products made of compacted thermally exfoliated graphite and products made of a composite material consisting of 100 layers of thin graphite foil and 99 layers of polyethylene, the differences in temperature fields did not exceed 1%. The obtained result substantiates our proposed approach to measuring thermal conductivity of flexible graphite foil by creating a multi-layer composite material. The thermal conductivity coefficient of such a composite at room temperature was <λGF> = 184 ± 6 W/m·K, which aligns well with measurements by the laser flash method. Full article
(This article belongs to the Section Carbon Materials)
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<p>Samples for measuring thermal conductivity coefficient.</p>
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<p>Distribution of the temperature (<b>a</b>) and thermal conductivity (<b>b</b>).</p>
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<p>Sample of graphite foil-based composite material. Artificial graphite sample G<sub>1</sub> (<b>a</b>), natural graphite sample G<sub>2</sub> (<b>b</b>), graphite foil-based composite material (<b>c</b>), geometric interpretation of the GF sample for modeling purposes (<b>d</b>).</p>
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<p>Photograph of the setup for measuring thermal conductivity of functional and structural products.</p>
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<p>Block diagram of the setup control.</p>
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<p>Geometric model of the setup (<b>a</b>) and temperature field distribution in the setup (<b>b</b>).</p>
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<p>Temperature distribution in the setup (<b>a</b>) and thermal conductivity coefficients (<b>b</b>) of samples G<sub>1</sub> и G<sub>2</sub>.</p>
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<p>Temperature distribution in the setup (<b>a</b>) and thermal conductivity coefficient of the GF sample (<b>b</b>).</p>
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<p>X-ray diffraction spectra of samples.</p>
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<p>Results of modeling the solid graphite foil sample: time to steady state (<b>a</b>) and temperature distribution along the setup (<b>b</b>).</p>
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16 pages, 2661 KiB  
Article
Influence of Mountain Wildfires on the Insulation Properties of Air Gaps in Power Grids
by Fangrong Zhou, Hao Geng, Gang Wen, Yutang Ma, Yi Ma, Guofang Wang, Jun Cao, Jiaze Xu and Hongwei Mei
Energies 2025, 18(2), 225; https://doi.org/10.3390/en18020225 - 7 Jan 2025
Viewed by 452
Abstract
The complex terrain of China frequently leads to wildfires, which in turn pose a threat to the safe operation of power transmission lines. Studying the breakdown characteristics of air gaps under wildfire conditions is of great significance for understanding wildfire propagation mechanisms, risk [...] Read more.
The complex terrain of China frequently leads to wildfires, which in turn pose a threat to the safe operation of power transmission lines. Studying the breakdown characteristics of air gaps under wildfire conditions is of great significance for understanding wildfire propagation mechanisms, risk assessment and management, and ecological environment protection. This paper establishes an experimental platform simulating wildfire climatic conditions and conducts experimental research on air gaps between rod–rod gaps and conductor–ground gaps. The experimental voltage types include direct current, power frequency, and standard operating waves. The impact of wildfire factors on the breakdown voltage and discharge characteristics of air gaps was obtained. The results indicate that the main factors affecting the air gap breakdown characteristics during wildfires are flame height and smoke. Flame height directly influences the gap insulation distance. Under flame bridging conditions, the maximum decrease in breakdown voltage reaches 70–80%. As the concentration of smoke increases, the degradation of insulation performance becomes more pronounced, with a reduction ranging from 20% to over 50%. Full article
(This article belongs to the Section F6: High Voltage)
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<p>Arrangement of ultra-high voltage outdoor test field: (<b>a</b>) the layout of the apparatus; (<b>b</b>) 7200 kV/720 kJ impulse voltage generator; (<b>c</b>) ±1600 kV/50 mA DC voltage generator; (<b>d</b>) 2250 kV power frequency voltage generator.</p>
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<p>Simulated wildfire test apparatus.</p>
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<p>The synchronization triggering process of the CCD high-speed camera.</p>
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<p>Wildfire test site.</p>
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<p>States of flame height: (<b>a</b>) flame bridging; (<b>b</b>) basic flame bridging; (<b>c</b>) flame height more than half; (<b>d</b>) visible flame.</p>
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<p>Conductor−ground wildfire tests under different fire conditions: (<b>a</b>) vertical fire intensity; (<b>b</b>) fire intensity experiencing wind deflection.</p>
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<p>Rod−rod air gap smoke test: (<b>a</b>) positive polarity DC; (<b>b</b>) negative polarity DC; (<b>c</b>) power frequency voltage.</p>
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15 pages, 3524 KiB  
Perspective
Electric Discharge-Generating Devices Developed for Pathogen, Insect Pest, and Weed Management: Current Status and Future Directions
by Shin-ichi Kusakari and Hideyoshi Toyoda
Agronomy 2025, 15(1), 123; https://doi.org/10.3390/agronomy15010123 - 6 Jan 2025
Viewed by 514
Abstract
Electrostatic techniques have introduced innovative approaches to devise efficient tools for pest control across various categories, encompassing pathogens, insects, and weeds. The focus on electric discharge technology has proven pivotal in establishing effective methods with simple device structures, enabling cost-effective fabrication using readily [...] Read more.
Electrostatic techniques have introduced innovative approaches to devise efficient tools for pest control across various categories, encompassing pathogens, insects, and weeds. The focus on electric discharge technology has proven pivotal in establishing effective methods with simple device structures, enabling cost-effective fabrication using readily available materials. The electric discharge-generating devices can be assembled using commonplace conductor materials, such as ordinary metal nets linked to a voltage booster and a grounded electric wire. The strategic pairing of charged and grounded conductors at specific intervals generates an electric field, leading the charged conductor to initiate a corona discharge in the surrounding space. As the applied voltage increases, the corona discharge intensifies and may eventually result in an arc discharge due to the breakdown of air when the voltage surpasses the insulation resistance limit. The utilization of corona and arc discharges plays a crucial role in these techniques, with the corona-discharging stage creating (1) negative ions to stick to pests, which can then be captured with a positively charged pole, (2) ozone gas to sterilize plant hydroponic solutions, and (3) plasma streams to exterminate fungal colonies on leaves, and the arc-discharging stage projecting electric sparks to zap and kill pests. These electric discharge phenomena have been harnessed to develop reliable devices capable of managing pests across diverse classes. In this review, we elucidate past achievements and challenges in device development, providing insights into the current status of research. Additionally, we discuss the future directions of research in this field, outlining potential avenues for further exploration and improvement. Full article
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<p>Schematic representation of the dual control system of rhizosphere and aerial pathogens for hydroponic plants by use of an ozone-generative spore precipitator [<a href="#B14-agronomy-15-00123" class="html-bibr">14</a>]. The device possessed multiple ozone-generative spore precipitation cylinders installed on a rectangular frame at a constant interval. The ozone produced in each cylinder was transferred to an ozone collection tank by an aspirator pump and supplied to the culture solution tank after adjusting the ozone concentration. Finally, the ozonized culture solution was circulated between the tank and hydroponic culture trough. Ozone is highly insoluble in water under normal temperature and pressure and can easily be converted into oxygen, such that it is impossible to dissolve ozone in water at the required concentrations for disinfection. Sterilization of pathogens in water was conducted by bubbling ozone-containing air into water. The strong oxidizing ability of ozone sterilizes microbes in the water when they come into contact with the air bubbles containing ozone. Black and blue arrows represent the direction of ozone gas and ozonized culture solution, respectively.</p>
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<p>(<b>A</b>) Schematic representation of the original spore-precipitation cylinder to trap airborne fungal spores [<a href="#B24-agronomy-15-00123" class="html-bibr">24</a>]. Copper wire (conductor) was covered with a transparent acrylic cylinder (insulator) and held at the middle, and both ends of the cylinder with insulating silicon stoppers. The cylinders were arranged in parallel. The copper wires were linked to a negative voltage booster (upper). The polarized dielectric cylinder produces an electrostatic field to cause the dipole on the spore. Spores were attracted to the cylinder through dielectrophoretic movement (lower). (<b>B</b>) Illustration depicting a spore-precipitation cylinder designed for ozone production [<a href="#B14-agronomy-15-00123" class="html-bibr">14</a>]. Each cylinder has a region at one end to generate ozone, and the remaining part of the cylinder is used to capture airborne spores. The copper wire was positively charged, and the electrostatic field was formed in the space around the dielectrically polarized cylinder. The ability to capture spores was the same as that of the original spore-precipitation cylinder. One tip end of the copper wire inside the acrylic cylinder was sharpened to create a needle pole. Continuous corona discharge (streamer discharge) was induced by the electric field between the tip end of the copper wire and a grounded copper ring attached to the cylinder edge. The ozone produced in each cylinder was transferred to a collection tank by an aspirator pump.</p>
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<p>(<b>A</b>) A portable plasma-stream exposer with a battery-operated voltage booster [<a href="#B26-agronomy-15-00123" class="html-bibr">26</a>]. (<b>B</b>) The diagram shows the spatial relationship between the charged probe tip and the grounded tomato leaf. The distance between the tip and the leaf was adjusted to determine the optimal tip position at which conidiophores were instantly destroyed by the plasma stream.</p>
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<p>Illustration of two devices designed for trapping tobacco smoke (<b>A</b>) [<a href="#B7-agronomy-15-00123" class="html-bibr">7</a>] and viruses enclosed in droplets (<b>B</b>) [<a href="#B6-agronomy-15-00123" class="html-bibr">6</a>] using corona discharge. The smoke-trapping device consisted of a spiked perforated metal plate (S-PMP) connected to a negative voltage booster (N-VB) and a grounded metal net (G-MN). On the other hand, the droplet-trapping device included the S-PMP connected to the N-VB, a water vessel (WV), and a plastic quadrangular hood (QH) with four legs. An electric field (EF) was formed between the spike tips and either the G-MN or the surface of the grounded water (G-W). The arrow indicates the direction of the ionic wind generated within the electric field.</p>
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<p>(<b>A</b>) Schematic representation of insect-mediated transient arc discharge generating device [<a href="#B16-agronomy-15-00123" class="html-bibr">16</a>]. The device consists of two identical metal nets: one was a negatively charged metal net linked to a negative voltage booster, and the other was a grounded metal net. The nets are arranged in parallel at a set interval, inducing corona discharge. An insulator board is attached to the outer surfaces of metal nets to prevent grains from falling through the net. (<b>B</b>) The glow corona is detected at the convex surfaces of the net, which act as needle poles. (<b>C</b>) When the insects enter the electric field between the nets at any location, the insects effectively become intermediate poles and are subjected to arc discharge-mediated sparks from the negatively charged metal net due to their conductive cuticle outer layer. Eventually, the electricity is transferred to the insect and then to the grounded net via a two-step arc discharge. Blue and red arrows represent corona and arc discharge, respectively.</p>
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<p>(<b>A</b>) Schematic representation of an electric soil cover designed to eliminate adult houseflies emerging from underground pupae [<a href="#B20-agronomy-15-00123" class="html-bibr">20</a>]. The cover consists of a negatively charged, non-insulated metal net (NC-MN) linked to a pulse-charging negative voltage booster (PN-VB) and a grounded metal net (G-MN). It is positioned on a plastic grating (PG). (<b>B</b>) Adult houseflies that emerge from underground pupae climb along the cell wall (green arrow) of the grating and encounter an arc discharge (red arrow) from the pulse-charged net when they reach the arcing zone. Subsequently, they are knocked downward (orange arrow) by the strong impact of the arc discharge.</p>
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<p>Schematic representation of the experimental devices used to eradicate weed seedlings emerging from the ground (cross-sectional view) (<b>A</b>,<b>B</b>) and kudzu vines climbing along the net (<b>C</b>). (<b>A</b>): The device consisted of a pulse-charged metal net (PC-MN) and a grounded metal net (G-MN) to emit arc-discharged mediated sparks to weed seedlings that have just emerged from ground soil [<a href="#B21-agronomy-15-00123" class="html-bibr">21</a>]. (<b>B</b>): The device possessed a single PC-MN with longer legs to emit sparks to the upper part of the seedling [<a href="#B22-agronomy-15-00123" class="html-bibr">22</a>]. (<b>C</b>): The device consisted of a single pulse-charged metal wire (PC-MW) and three grounded metal wires (G-MN1-3) attached to a polypropylene net (PN). The red arrow represents the arc discharge from the charged net and wire [<a href="#B44-agronomy-15-00123" class="html-bibr">44</a>].</p>
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<p>(<b>A</b>–<b>C</b>) Images depicting the successful application of pulse-charged metal nets for controlling weed seedlings on a slope [<a href="#B22-agronomy-15-00123" class="html-bibr">22</a>]. (<b>A</b>) shows the setup at the beginning of the experiment, while (<b>B</b>) displays the condition after 3 months. (<b>C</b>), revealing thriving weed seedlings beneath the nets with none penetrating through. (<b>D</b>,<b>E</b>) (<b>E</b>) shows a pulse-charged metal wire (PC-MW) and a grounded metal wire (G-MW) affixed to a fence with nearby kudzu plants for testing purposes, showing fence-climbing vines with their apical tips damaged due to exposure to arc discharge [<a href="#B44-agronomy-15-00123" class="html-bibr">44</a>]. Image E provides a close-up of (<b>D</b>), revealing damaged apical tips.</p>
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