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Search Results (1,133)

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13 pages, 12886 KiB  
Article
Investigation of the Microstructures and Mechanical Properties of Sn-Cu-Bi-In-Ni Solders
by Xiaochun Lv, Chenghao Zhang, Yang Liu, Zhen Pan, Zhiyuan Wang and Fenglian Sun
Materials 2025, 18(4), 858; https://doi.org/10.3390/ma18040858 (registering DOI) - 16 Feb 2025
Abstract
The development of Ag-free Sn solders has attracted significant attention due to the requirement of high-density electronic packaging. In this study, we investigate the Ni element on the microstructures and mechanical properties of Ag-free Sn-Cu-Bi-In solders. This paper details the microstructures and phases [...] Read more.
The development of Ag-free Sn solders has attracted significant attention due to the requirement of high-density electronic packaging. In this study, we investigate the Ni element on the microstructures and mechanical properties of Ag-free Sn-Cu-Bi-In solders. This paper details the microstructures and phases of the as-prepared Sn-Cu-Bi-In-Ni solders, as well as its mechanical properties. Specifically, the intermetallic compound (IMC) Cu6Sn5 is observed to be distributed in the Sn matrix, forming near-eutectic structures. The incorporation of Ni into Sn-Cu-Bi-In enhances the mechanical properties of the solder joints, including the shear strength and vibrational stability. In the joint obtained using the as-prepared Sn-Cu-Bi-In-Ni solders, a (Cu,Ni)6Sn5 IMC layer forms at the interface between Sn ball and Cu pad. The beneficial effects of Ni can be primarily attributed to its ability to adjust the mechanical properties and thermal expansion, enhancing the stability of solder joints. A TEM analysis reveals the closely packed atomic interface of Cu/(Cu,Ni)6Sn5 and (Cu,Ni)6Sn5/Sn, elucidating the joining mechanism involved. Full article
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Figure 1
<p>Schematic diagram of preparing Sn-Cu-Bi-In-Ni solder alloys, reflow soldering, shear test, and vibration test.</p>
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<p>Characterization of as-prepared Sn-Cu-Bi-In-Ni solders: (<b>a</b>) SEM images, element map distribution of (<b>b</b>) Sn, (<b>c</b>) Cu, (<b>d</b>) Bi, (<b>e</b>) In, and (<b>f</b>) Ni; (<b>g</b>) EDS spectra of the whole region in (<b>a</b>); (<b>h</b>) DSC results of Sn-Cu-Bi-In-Ni and SAC305 solders.</p>
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<p>(<b>a</b>) STEM image of Sn-Cu-Bi-In-Ni solders, corresponding element map distribution of (<b>b</b>) Sn, (<b>c</b>) Cu, (<b>d</b>) In, (<b>e</b>) Bi, and (<b>f</b>) Ni.</p>
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<p>(<b>a</b>) Bright-field TEM image of Sn-Cu-Bi-In-Ni solder, HRTEM images of (<b>b</b>) region A and (<b>c</b>) region B in (<b>a</b>), (<b>d</b>) magnified image of selected region in (<b>c</b>) (the inset image is the FFT pattern of the selected region in (<b>c</b>), and the circles are for pointing the diffraction spots out), atomic spacing in the selected region of (<b>b</b>): (<b>e</b>,<b>d</b>): (<b>f</b>).</p>
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<p>(<b>a</b>) Tensile strength of SAC305, Sn-Cu-Bi-In and Sn-Cu-Bi-In-Ni solders, (<b>b</b>) fracture morphology of Sn-Cu-Bi-In-Ni solders, (<b>c</b>) XRD patterns of the Sn-Cu-Bi-In-Ni solders.</p>
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<p>(<b>a</b>) Cross-sectional SEM image of the Sn-Cu-Bi-In-Ni solder joints on Cu pads (the inset table is the EDS result of Point A); (<b>b</b>) shear strength of the joints prepared by three kinds of solders.</p>
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<p>(<b>a</b>) STEM image of the Sn-Cu-Bi-In-Ni solder joint, corresponding element map distribution of (<b>b</b>) Sn, (<b>c</b>) Cu, (<b>d</b>) Bi, (<b>e</b>) In, and (<b>f</b>) Ni.</p>
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<p>TEM analysis of the Sn-Cu-Bi-In-Ni solder joint. (<b>a</b>) Bright-field TEM image, SAED patterns of (<b>b</b>) Cu substrate, (<b>c</b>) (Cu,Ni)-Sn compound, and (<b>d</b>) Sn ball; HRTEM images of (<b>e</b>) (Cu,Ni)-Sn compound and (<b>f</b>) Sn ball; interfacial regions of (<b>g</b>) Cu/(Cu,Ni)<sub>6</sub>Sn<sub>5</sub> and (<b>h</b>) Sn/(Cu,Ni)<sub>6</sub>Sn<sub>5</sub>.</p>
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<p>(<b>a</b>) SEM image of the fracture of Sn-Cu-Bi-In-Ni joints, corresponding element map distribution of (<b>b</b>) Sn, (<b>c</b>) Cu, (<b>d</b>) In, (<b>e</b>) Bi, and (<b>f</b>) Ni.</p>
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<p>Vibration test of the Sn-Cu-Bi-In-Ni solder joints: (<b>a</b>) schematic diagram of vibration platform, (<b>b</b>) SEM image of the interfacial cracks after vibration test, (<b>c</b>) vibrational lifetime of different solder joints.</p>
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11 pages, 2464 KiB  
Article
Computer Simulation of the Natural Vibrations of a Rigidly Fixed Plate Considering Temperature Shock
by Andry Sedelnikov, Sergey Glushkov, Maksim Evtushenko, Yurii Skvortsov and Alexandra Nikolaeva
Computation 2025, 13(2), 49; https://doi.org/10.3390/computation13020049 - 10 Feb 2025
Abstract
This paper presents the results of a computational experiment on the natural vibrations of a homogeneous rigidly fixed plate after a temperature shock. Unlike in many well-known studies, in this work, the plate is not stationary at the moment of thermal shock. This [...] Read more.
This paper presents the results of a computational experiment on the natural vibrations of a homogeneous rigidly fixed plate after a temperature shock. Unlike in many well-known studies, in this work, the plate is not stationary at the moment of thermal shock. This formulation has wide practical applications. For example, as a result of the unfolding of solar panels, free vibrations are excited. The purpose of this work was to analyze the effect of temperature shock on the characteristics of the plate’s own vibrations. Specifying the parameters of natural vibrations and considering temperature shock make it possible to model the vibration process more adequately. The simulation parameters simulate the conditions of the space environment. Therefore, the results of this study can be applied to the study of thermal vibrations in solar panels and other large elastic elements of spacecraft. Full article
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<p>The appearance of the simulated plate with a grid of elements in the ANSYS environment (<b>a</b>) and its enlarged fragment (<b>b</b>).</p>
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<p>The appearance of the rigid fixing of the plate edge.</p>
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<p>Natural vibrations of the free edge of the plate, resulting from its initial deformation.</p>
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<p>The difference between the numerical values of the natural frequencies and the approximate theoretical frequency (1), depending on the size of the mesh.</p>
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<p>Plate temperature dynamics: 1—surface layer exposed to temperature shock; 2—shadow surface layer.</p>
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<p>Natural vibrations of the free edge of the plate, considering temperature shock.</p>
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13 pages, 525 KiB  
Article
The Influence of Mass on Dynamic Response of Cracked Timoshenko Beam with Restrained End Conditions: The Truncated Theory
by Maria Anna De Rosa, Carla Ceraldi, Hector D. Martin, Antonella Onorato, Marcelo Tulio Piovan and Maria Lippiello
Appl. Mech. 2025, 6(1), 11; https://doi.org/10.3390/applmech6010011 - 7 Feb 2025
Abstract
In this paper, the dynamic response of the Timoshenko cracked beam subjected to a mass is investigated. In turn, it is assumed that the beam has its ends restrained with both transverse and rotational elastic springs. Based on an alternative beam theory, truncated [...] Read more.
In this paper, the dynamic response of the Timoshenko cracked beam subjected to a mass is investigated. In turn, it is assumed that the beam has its ends restrained with both transverse and rotational elastic springs. Based on an alternative beam theory, truncated Timoshenko theory (TTT), the governing equations of motion of the cracked beam are derived and the influence of a mass on the behavior of free vibrations is investigated. The novelty of the proposed approach lies in the fact that the variational method used in the truncated theory simplifies the derivation of the equation of motion via the classical theory, and the perfect analogy between the two theories is shown. The objective of the present formulation lies in finding the equations of the truncated Timoshenko model with their corresponding boundary conditions and establishing their mathematical similarity with the geometric approach. It is shown that the differential equations with their corresponding boundary conditions, used to solve the dynamic problem of Timoshenko truncated beams through variational formulations, have the same form as those obtained through the direct method. Finally, some numerical examples are carried out to evaluate the influence of a mass and its position on the vibration performances of the cracked Timoshenko model. Additionally, the effects of the crack positions, the shear deformation and rotational inertia, and the yielding constraints on the natural frequencies are also discussed in some numerical examples. Full article
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<p>Beam model under consideration.</p>
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<p>First three dimensionless frequencies as rotational stiffness varies.</p>
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<p>First three dimensionless frequencies as translational rigidity varies.</p>
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<p>First dimensionless frequency <math display="inline"><semantics> <msub> <mo>Ω</mo> <mn>1</mn> </msub> </semantics></math> for three different values of mass and crack positions.</p>
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<p>Effect of crack and mass at different locations on Mode-I: (<b>a</b>) <math display="inline"><semantics> <msub> <mi>λ</mi> <mi>M</mi> </msub> </semantics></math> = 0.3 and (<b>b</b>) <math display="inline"><semantics> <msub> <mi>λ</mi> <mi>M</mi> </msub> </semantics></math> = 0.7.</p>
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29 pages, 43807 KiB  
Article
Insights on the Influence of the Central-Cut Width of the Box Assembly with Removable Component on Its Dynamical Responses
by Christopher Padilla, Antonio Flores, Jonah Madrid, Ezekiel Granillo and Abdessattar Abdelkefi
Appl. Sci. 2025, 15(3), 1671; https://doi.org/10.3390/app15031671 - 6 Feb 2025
Abstract
This investigation focuses on the dynamical effects caused by varying the central-cut width within the Box Assembly with Removable Component (BARC) system. The central-cut widths included in this study are a 0.5″ cut, a 0.25″ cut, a thin 0.1″ cut, and a structure [...] Read more.
This investigation focuses on the dynamical effects caused by varying the central-cut width within the Box Assembly with Removable Component (BARC) system. The central-cut widths included in this study are a 0.5″ cut, a 0.25″ cut, a thin 0.1″ cut, and a structure that did not have a cut at all. Finite element analysis was conducted to determine the mode shapes and natural frequencies of each of the BARC structures. Structural dynamics experiments were run to examine the effects of the central-cut width on the dynamical responses and nonlinear characteristics of the BARC system. Free vibration testing with an impact hammer was carried out to excite the system and extract the dominant frequencies and directions of the significant responses. A pseudorandom vibration test that allows for the qualitative determination of any nonlinear behavior within the system was performed. This type of behavior can include nonlinear softening, nonlinear hardening, and the most common, nonlinear damping due to the presence of several bolted-joint connections and the possible activation of geometric and inertia nonlinearities. To quantitatively investigate the impacts of the central-cut width on the dynamics of the system, swept sinusoidal testing was conducted. It is determined that almost all systems with central cuts demonstrate the presence of nonlinear softening, but at times, nonlinear hardening trends are seen, particularly in the 0.1″ cut and no-cut systems when testing harmonically. Each of the central-cut systems displays nonlinear damping, with the amount of damping generally increasing as the central cut decreases in size. The effect of the central cut of the BARC system on the mode-switching ability of the system is negligible; however, mode switching takes place when comparing the central-cut configurations to the no-cut one. These results show the significance of accurately measuring the central-cut width and how geometric uncertainty may change the dynamical responses and nonlinear properties of the system. Full article
(This article belongs to the Special Issue Vibration Problems in Engineering Science)
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<p>Computer-aided design (CAD) models of each BARC system: (<b>a</b>) 0.5″ cut, (<b>b</b>) 0.25″ cut, (<b>c</b>) 0.1″ cut, and (<b>d</b>) no cut.</p>
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<p>Accelerometers’ locations on the BARC system.</p>
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<p>Alternating screw pattern for the removable component.</p>
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<p>BARC base hole pattern.</p>
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<p>The impact testing setup for the BARC system.</p>
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<p>Random vibration and harmonic vibration testing setup utilizing a slip-table setup for testing of the BARC system.</p>
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<p>Impact testing results from the side accelerometer: (<b>a</b>) 0.5″ cut, (<b>b</b>) 0.25″ cut, (<b>c</b>) 0.1″ cut, and (<b>d</b>) no-cut BARC systems.</p>
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<p>Impact testing results from the side accelerometer: (<b>a</b>) 0.5″ cut, (<b>b</b>) 0.25″ cut, (<b>c</b>) 0.1″ cut, and (<b>d</b>) no-cut BARC systems.</p>
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<p>Impact testing results from the RC accelerometer: (<b>a</b>) 0.5″ cut, (<b>b</b>) 0.25″ cut, (<b>c</b>) 0.1″ cut, and (<b>d</b>) no-cut BARC systems.</p>
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<p>Impact testing results from the RC accelerometer: (<b>a</b>) 0.5″ cut, (<b>b</b>) 0.25″ cut, (<b>c</b>) 0.1″ cut, and (<b>d</b>) no-cut BARC systems.</p>
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<p>Impact vibration testing comparing (<b>a</b>) X-direction, (<b>b</b>) Y-direction, and (<b>c</b>) Z-direction responses from different accelerometers for the 0.5″ cut BARC system.</p>
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<p>A comparison of impact results using (<b>a</b>) the side accelerometer and (<b>b</b>) the RC accelerometer.</p>
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<p>Random vibration results from the Y-direction of the side accelerometer: (<b>a</b>) 0.5″ cut, (<b>b</b>) 0.25″ cut, (<b>c</b>) 0.1″ cut, and (<b>d</b>) no-cut BARC systems.</p>
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<p>Random vibration results from the Y-direction of the side accelerometer: (<b>a</b>) 0.5″ cut, (<b>b</b>) 0.25″ cut, (<b>c</b>) 0.1″ cut, and (<b>d</b>) no-cut BARC systems.</p>
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<p>Random vibration results from the RC accelerometer in the Y-direction: (<b>a</b>) 0.5″ cut, (<b>b</b>) 0.25″ cut, (<b>c</b>) 0.1″ cut, and (<b>d</b>) no-cut BARC systems.</p>
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<p>Random vibration results using the RC accelerometer in the Z-direction: (<b>a</b>) 0.5″ cut, (<b>b</b>) 0.25″ cut, (<b>c</b>) 0.1″ cut, and (<b>d</b>) no-cut BARC systems.</p>
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<p>A 3-D comparison of all cut widths for the (<b>a</b>) side and (<b>b</b>) RC accelerometers using data from the 1 × 10<sup>−4</sup> V<sup>2</sup>/Hz excitation.</p>
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<p>A comparison of BARC central-cut width configurations using random vibration testing from the (<b>a</b>) side Y, (<b>b</b>) RC Y, and (<b>c</b>) RC Z accelerometers at 1 × 10<sup>−4</sup> V<sup>2</sup>/Hz excitation.</p>
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<p>Harmonic vibration results from the Y-direction of the side accelerometer: (<b>a</b>) 0.5″ cut, (<b>b</b>) 0.25″ cut, (<b>c</b>) 0.1″ cut, and (<b>d</b>) no-cut BARC systems.</p>
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<p>A comparison of central-cut width configurations using the side accelerometer in the Y-direction for the 50 mV and 150 mV input excitation levels.</p>
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<p>Harmonic vibration results from the RC Y-direction accelerometer from the (<b>a</b>) 0.5″ cut, (<b>b</b>) 0.25″ cut, (<b>c</b>) 0.1″ cut, and (<b>d</b>) no-cut BARC systems.</p>
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<p>A comparison of central-cut width configurations using the RC accelerometer in the Y-direction for the 50 mV and 150 mV input levels of excitation.</p>
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<p>A comparison using the Y-direction data from the side and RC accelerometers and Z-direction data from the RC accelerometer using the (<b>a</b>) 50 mV and (<b>b</b>) 150 mV harmonic excitation data.</p>
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<p>Time histories found around resonance from harmonic testing for the 0.5″ cut configuration with 150 mV excitation for the (<b>a</b>) side accelerometer and (<b>b</b>) RC accelerometer in the Y-direction and (<b>c</b>) RC accelerometer in the Z-direction.</p>
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<p>Q-Factor plot for the RC accelerometer in the Y-direction.</p>
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17 pages, 4722 KiB  
Article
Research on Space Maglev Vibration Isolation Control System Modeling and Simulation
by Mao Ye and Jianyu Wang
Appl. Sci. 2025, 15(3), 1648; https://doi.org/10.3390/app15031648 - 6 Feb 2025
Abstract
The working accuracy of space optical payloads and sensitive components carried on space aircraft greatly depends on the pointing accuracy and stability of the platform. Based on Disturbance Free Payload (DFP) technology, non-contact maglev technology is proposed in this paper, achieving dynamic and [...] Read more.
The working accuracy of space optical payloads and sensitive components carried on space aircraft greatly depends on the pointing accuracy and stability of the platform. Based on Disturbance Free Payload (DFP) technology, non-contact maglev technology is proposed in this paper, achieving dynamic and static isolation of the platform module and payload module, so that the vibration and interference of the platform module with movable and flexible components will not be transmitted to the payload module, thereby achieving the effect of vibration isolation. High-precision active control of the payload module is adopted at the same time; the platform module follows the master–slave collaborative control strategy of the payload module, meeting the requirements of high-performance payloads. A primary and backup redundant controller is designed, using a one-to-four architecture. The control board achieves high-speed and high-precision driving current control, voltage output, and outputs current feedback signal sampling. Based on uniform magnetic field design, high-precision force control performance is ensured by adjusting current accuracy. Interdisciplinary joint simulation of electric, magnetic, and structural aspects was conducted on the magnetic levitation isolation system. By conducting physical testing and calibration and designing a testing and calibration system, it has been proven that the system meets the design requirements, achieving high-precision current control technology of 0.15 mA and driving force control technology of 0.5 mN. Full article
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<p>The master–slave collaborative control strategy.</p>
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<p>Overall scheme.</p>
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<p>Actuator structure.</p>
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<p>Control principle block diagram.</p>
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<p>Current control principle diagram.</p>
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<p>Bidirectional four-channel coil current power board.</p>
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<p>Magnetic simulation.</p>
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<p>Maglev modeling.</p>
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<p>Three-dimensional model of the actuator.</p>
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<p>The magnetic density cloud maps and vector diagrams of XY plane.</p>
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<p>The magnetic induction intensity.</p>
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<p>The variation curves of current.</p>
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<p>The variation curves of electromagnetic force.</p>
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<p>Comprehensive performance testing calibration process.</p>
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<p>Actuator prototype sample.</p>
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17 pages, 34339 KiB  
Article
Prediction and Optimization of the Long-Term Fatigue Life of a Composite Hydrogen Storage Vessel Under Random Vibration
by Xiaoshuang Xiong, Wentao Wang, Xiang Li, Fei Fan, Jiacheng Zhou and Mingzhang Chen
Materials 2025, 18(3), 712; https://doi.org/10.3390/ma18030712 - 6 Feb 2025
Abstract
A composite hydrogen storage vessel (CHSV) is one key component of the hydrogen fuel cell vehicle, which always suffers random vibration during transportation, resulting in fatigue failure and a reduction in service life. In this paper, firstly, the free and constrained modes of [...] Read more.
A composite hydrogen storage vessel (CHSV) is one key component of the hydrogen fuel cell vehicle, which always suffers random vibration during transportation, resulting in fatigue failure and a reduction in service life. In this paper, firstly, the free and constrained modes of CHSV are experimentally studied and numerically simulated. Subsequently, the random vibration simulation of CHSV is carried out to predict the stress distribution, while Steinberg’s method and Dirlik’s method are used to predict the fatigue life of CHSV based on the results of stress distribution. In the end, the optimization of ply parameters of the composite winding layer was conducted to improve the stress distribution and fatigue life of CHSV. The results show that the vibration pattern and frequency of the free and constrained modes of CHSV obtained from the experiment tests and the numerical predictions show a good agreement. The maximum difference in the value of the vibration frequency of the free and constrained modes of CHSV from the FEA and experiment tests are, respectively, 8.9% and 8.0%, verifying the accuracy of the finite element model of CHSV. There is no obvious difference between the fatigue life of the winding layer and the inner liner calculated by Steinberg’s method and Dirlik’s method, indicating the accuracy of FEA of fatigue life in the software Fe-safe. Without the optimization, the maximum stresses of the winding layer and the inner liner are found to be near the head section by 469.4 MPa and 173.0 MPa, respectively, and the numbers of life cycles of the winding layer and the inner liner obtained based on the Dirlik’s method are around 1.66 × 106 and 3.06 × 106, respectively. Through the optimization of ply parameters of the composite winding layer, the maximum stresses of the winding layer and the inner liner are reduced by 66% and 85%, respectively, while the numbers of life cycles of the winding layer and the inner liner both are increased to 1 × 107 (high cycle fatigue life standard). The results of the study provide theoretical guidance for the design and optimization of CHSV under random vibration. Full article
(This article belongs to the Special Issue Advances in Modelling and Simulation of Materials in Applied Sciences)
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<p>Geometric structure diagram of CHSV.</p>
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<p>Schematic diagram of (<b>A</b>) the free modal experiment and (<b>B</b>) the constrained modal experiment of CHSV.</p>
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<p>Finite element model of CHSV.</p>
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<p>Finite element model and boundary condition of constrained mode and random vibration simulation.</p>
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<p>Acceleration power spectral density [<a href="#B34-materials-18-00712" class="html-bibr">34</a>].</p>
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<p>The first four orders of free modes of CHSV: (<b>a</b>) the finite element prediction of the vibration modes, (<b>b</b>) the experimental result of the vibration modes, and (<b>c</b>) the comparison of the vibration frequencies.</p>
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<p>The first four-orders of constrained modes of CHSV: (<b>a</b>) the finite element prediction of vibration modes, (<b>b</b>) the experimental result of vibration modes, and (<b>c</b>) the comparison of vibration frequencies.</p>
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<p>Stress of the (<b>a</b>) winding layer and (<b>b</b>) inner liner under random vibration.</p>
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<p>Rmises stress of the (<b>a</b>) winding layer and (<b>b</b>) inner liner under random vibration.</p>
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<p>Maximum vibration deformation of (<b>a</b>) the winding layer and (<b>b</b>) the inner liner of the CHSV under random vibration.</p>
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<p>Fatigue life curve of aluminum alloy Al6061-T6 [<a href="#B36-materials-18-00712" class="html-bibr">36</a>].</p>
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<p>Fatigue life curve of CFRP [<a href="#B37-materials-18-00712" class="html-bibr">37</a>].</p>
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<p>Fatigue life of (<b>a</b>) the winding layer and (<b>b</b>) the inner liner of the CHSV using software Fe-safe.</p>
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<p>Comparison of winding layers: (<b>a</b>) before optimization and (<b>b</b>) after optimization.</p>
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<p>Stress of (<b>a</b>) the winding layer and (<b>b</b>) the inner liner of the CHSV after optimization.</p>
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<p>The fatigue life of (<b>a</b>) the winding layer and (<b>b</b>) the inner liner of the CHSV after optimization.</p>
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24 pages, 4555 KiB  
Review
Biophysics of Voice Onset: A Comprehensive Overview
by Philippe H. DeJonckere and Jean Lebacq
Bioengineering 2025, 12(2), 155; https://doi.org/10.3390/bioengineering12020155 - 6 Feb 2025
Abstract
Voice onset is the sequence of events between the first detectable movement of the vocal folds (VFs) and the stable vibration of the vocal folds. It is considered a critical phase of phonation, and the different modalities of voice onset and their distinctive [...] Read more.
Voice onset is the sequence of events between the first detectable movement of the vocal folds (VFs) and the stable vibration of the vocal folds. It is considered a critical phase of phonation, and the different modalities of voice onset and their distinctive characteristics are analysed. Oscillation of the VFs can start from either a closed glottis with no airflow or an open glottis with airflow. The objective of this article is to provide a comprehensive survey of this transient phenomenon, from a biomechanical point of view, in normal modal (i.e., nonpathological) conditions of vocal emission. This synthetic overview mainly relies upon a number of recent experimental studies, all based on in vivo physiological measurements, and using a common, original and consistent methodology which combines high-speed imaging, sound analysis, electro-, photo-, flow- and ultrasound glottography. In this way, the two basic parameters—the instantaneous glottal area and the airflow—can be measured, and the instantaneous intraglottal pressure can be automatically calculated from the combined records, which gives a detailed insight, both qualitative and quantitative, into the onset phenomenon. The similarity of the methodology enables a link to be made with the biomechanics of sustained phonation. Essential is the temporal relationship between the glottal area and intraglottal pressure. The three key findings are (1) From the initial onset cycles onwards, the intraglottal pressure signal leads that of the opening signal, as in sustained voicing, which is the basic condition for an energy transfer from the lung pressure to the VF tissue. (2) This phase lead is primarily due to the skewing of the airflow curve to the right with respect to the glottal area curve, a consequence of the compressibility of air and the inertance of the vocal tract. (3) In case of a soft, physiological onset, the glottis shows a spindle-shaped configuration just before the oscillation begins. Using the same parameters (airflow, glottal area, intraglottal pressure), the mechanism of triggering the oscillation can be explained by the intraglottal aerodynamic condition. From the first cycles on, the VFs oscillate on either side of a paramedian axis. The amplitude of these free oscillations increases progressively before the first contact on the midline. Whether the first movement is lateral or medial cannot be defined. Moreover, this comprehensive synthesis of onset biomechanics and the links it creates sheds new light on comparable phenomena at the level of sound attack in wind instruments, as well as phenomena such as the production of intervals in the sung voice. Full article
(This article belongs to the Special Issue The Biophysics of Vocal Onset)
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<p>VKG (single-line scan) at approximately the midpoint of VF length, with the related images obtained from high-speed video just before the onset of VF vibration. Total duration about 34 ms. Left: physiological onset starting from an open glottis; right: physiological onset starting from a closed glottis. Hence, there is a wide range of voice onset modalities, but in experimental (physiological) conditions (i.e., on normal subjects, trained vocalists and singers) most authors interested in voice physiology make the distinction between the ‘soft’ (albeit slightly breathy, for clarity) onset, <span class="underline">with</span> airflow, and the ‘hard’ onset <span class="underline">without</span> airflow.</p>
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<p>Physiological soft <sub>(o)</sub> onset: two consecutive images from Fink’s 16 frames/s laryngoscopic film: the picture on the left is still sharp, and points out the immobile spindle-shaped configuration, while in the picture on the right (taken 62.5 ms later) vibration has started, evidenced by blurring. Scale is not indicated on Fink’s film, but the length of the vibrating VFs must be about 15 mm.</p>
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<p>Snapshot extracted from a videolaryngoscopic recording (showing glottis and VFs) in a normal male trained vocalist uttering a physiological sustained /a:/ at 130 Hz (soft <sub>(o)</sub> onset). The time interval between the left and right images is 30 ms. The typical (sharp) pre-onset spindle-shaped glottal configuration appears clearly in the left image. In the right image, blurring indicates that VFs are vibrating. The length of the vibrating part is 13 mm.</p>
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<p>Idem as <a href="#bioengineering-12-00155-f003" class="html-fig">Figure 3</a>, with a different voicing condition (slightly higher and louder, 145 Hz), and a slightly different pre-onset glottal configuration.</p>
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<p>Two images preceding the oscillation by 30 ms, demonstrating how the pre-oscillatory size of the ellipsoid may differ to a large extent.</p>
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<p>Example of a physiological soft onset (125 Hz), but starting from a closed glottis (soft <sub>(c)</sub>): Left: the VFs are completely (but gently) adducted, without ‘effort closure’ that should tend to-wards a sphincteric behaviour. Right: (next frame, 30 ms later) the VF edges appear blurred, similarly to the soft <sub>(o)</sub> examples.</p>
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<p>Two examples of pre-onset configurations that represent a (moderate) exaggeration of the soft <sub>(o)</sub> and soft <sub>(c)</sub> patterns, respectively: the picture on the left is the start position of a breathy onset, with a significantly larger glottal area, and the picture on the right is the start position of a clearly hard onset, with adduction of the ventricular folds.</p>
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<p>VKG at four equidistant levels of the vibrating glottis, obtained from high-speed video (2000 frames/s). Soft <sub>(o)</sub>, slightly breathy onset. Time is progressing from top to bottom. /a:/; healthy male vocalist (~125 Hz; 65 dB at 10 cm).</p>
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<p>Spindle-shaped glottal split just before the oscillation starts in a soft <sub>(o)</sub>, slightly breathy onset (snapshot from high-speed video 2000 frames/s). Healthy male vocalist (~125 Hz; 65 dB at 10 cm).</p>
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<p>Onset from a closed glottis. From top to bottom: FGG, EGG, PGG and sound oscillogram. Modal phonation (~125 Hz; 65 dB at 10 cm).</p>
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<p>Onset from an open glottis. From top to bottom: PGG (glottal area), ultrasound signal, EGG and FGG as a function of time. Comfortable pitch and loudness (~125 Hz; 65 dB at 10 cm). As can be seen on the glottal area and flow traces, a closed plateau first clearly appears after the fourth cycle. The EGG trace indicates that a very brief and limited contact already occurred in the previous cycle.</p>
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<p>Breathy onset. From top to bottom: FGG, EGG, PGG (raw tracings). Comfortable pitch and loudness (~125 Hz; 65 dB at 10 cm). Similar pattern to a soft <sub>(o)</sub> onset, but slower progression: the amplitude of oscillations progressively increases over more than 10 cycles.</p>
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<p>Soft <sub>(o)</sub> onset (~125 Hz; 65 dB at 10 cm) From top to bottom: PGG, FGG and intraglottal pressure. A phase lead (slightly less than 90°) of the intraglottal pressure with respect to the glottal opening is observed. Time in ms.</p>
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<p>Progress of phase lead (ms) of the intraglottal pressure with respect to the glottal area as a function of the sequential number of the cycle during the soft <sub>(o)</sub>/breathy onsets. The number of cases falls as the number of the cycle increases. The average phase lead of the pressure rises from 0 up to about 0.9 ms.</p>
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<p>Typical example of a soft <sub>(o)</sub> onset recording. From top to bottom: FGG (Rothenberg mask), EGG and light signal proportional to the glottal area (PGG). Total duration of the recording is 124 ms. On the area trace, the level at which oscillation starts and the maximal area amplitude (100%)—which can be calibrated on images—are indicated by vertical arrows. The flow level at which oscillation starts is indicated by the vertical arrow on the airflow trace with respect to the baseline (flow = 0) reached when complete glottal closure is observed.</p>
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<p>Equivalent diameter (mm) calculated from the measured glottal area as a function of the velocity of air particles (m/s) at the start of oscillation.</p>
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<p>Amplitude (y-axis) of glottal area oscillations during the first cycles (n = 35). Arbitrary units, linear scale. Cycle numbers are on the x-axis.</p>
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<p>Number of identifiable oscillation cycles before steady-state VF vibration is reached in three different signals: FGG, EGG and PGG.</p>
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<p>Example of the evolution of cycle duration over the first 18 cycles of a soft <sub>(o)</sub>/breathy onset (125 Hz; 65 dB at 10 cm). Cycle numbers are on the x-axis. The insert shows the glottal area trace. Slight progressive decrease in the fundamental frequency.</p>
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<p>Period duration of the first 10 cycles of a sustained /a:/ in a normal male subject (10 repetitions of a soft <sub>(o)</sub> voice onset). Period duration becomes stable after about 5 cycles.</p>
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<p>Pitch-matching task for a standardised interval. Principle of measurement: X-axis is time, Y-axis is fundamental frequency. The singer makes a pitch jump (third, fifth or octave) without legato, i.e., with a short interruption in vocal emission. Cycle duration is measured just before (10 cycles) and just after (5 and 10 cycles) the pitch jump.</p>
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<p>Lip configuration of a professional horn player, as filmed through a transparent mouthpiece. The picture precedes the vibration onset by 50 ms. The lips are slightly parted, and there is a small airflow escape.</p>
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<p>Attack of a sound of 175 Hz (horn). Simultaneous recording of EGG and sound oscillogram (SO). On the EGG signal, the maximum of VF contact is below, the minimum is above. Five glottal cycles precede the emission of sound. The amplitude of glottal movements is larger during these five first cycles than during the sustained emission. The glottal frequency of the first cycles is slightly higher than that of the sound to be played, but the adjustment is very quick.</p>
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<p>Vibratory double-reed of a tenor crumhorn, with its typical spindle shape. When the instrument is played, a cap covers this double reed, preventing any contact with the player’s lips.</p>
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15 pages, 6241 KiB  
Article
Modal Parameter Identification of the Improved Random Decrement Technique-Stochastic Subspace Identification Method Under Non-Stationary Excitation
by Jinzhi Wu, Jie Hu, Ming Ma, Chengfei Zhang, Zenan Ma, Chunjuan Zhou and Guojun Sun
Appl. Sci. 2025, 15(3), 1398; https://doi.org/10.3390/app15031398 - 29 Jan 2025
Abstract
Commonly used methods for identifying modal parameters under environmental excitations assume that the unknown environmental input is a stationary white noise sequence. For large-scale civil structures, actual environmental excitations, such as wind gusts and impact loads, cannot usually meet this condition, and exhibit [...] Read more.
Commonly used methods for identifying modal parameters under environmental excitations assume that the unknown environmental input is a stationary white noise sequence. For large-scale civil structures, actual environmental excitations, such as wind gusts and impact loads, cannot usually meet this condition, and exhibit obvious non-stationary and non-white-noise characteristics. The theoretical basis of the stochastic subspace method is the state-space equation in the time domain, while the state-space equation of the system is only applicable to linear systems. Therefore, under non-smooth excitation, this paper proposes a stochastic subspace method based on RDT. Firstly, this paper uses the random decrement technique of non-stationary excitation to obtain the free attenuation response of the response signal, and then uses the stochastic subspace identification (SSI) method to identify the modal parameters. This not only improves the signal-to-noise ratio of the signal, but also improves the computational efficiency significantly. A non-stationary excitation is applied to the spatial grid structure model, and the RDT-SSI method is used to identify the modal parameters. The identification results show that the proposed method can solve the problem of identifying structural modal parameters under non-stationary excitation. This method is applied to the actual health monitoring of stadium grids, and can also obtain better identification results in frequency, damping ratio, and vibration mode, while also significantly improving computational efficiency. Full article
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<p>Flowchart of identification.</p>
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<p>Rectangular slab grid structure.</p>
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<p>Theoretical frequencies and vibration shapes of the first six orders.</p>
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<p>Non-stationary white noise signal.</p>
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<p>Damping ratio identification results.</p>
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<p>Vibration mode identification results.</p>
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<p>Stadium east suspension grids.</p>
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<p>Layout of measuring points.</p>
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<p>Photograph of sensor arrangement.</p>
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<p>Acquisition equipment collecting data on-site.</p>
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<p>Theoretical frequencies and modes of vibration of the first 6 orders.</p>
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<p>Response signal at measurement point 4.</p>
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<p>Results of mode shape identification by the improved RDT-SSI method.</p>
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<p>Damping ratio identification results.</p>
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26 pages, 8754 KiB  
Article
Weight Effects on Vertical Transverse Vibration of a Beam with a Nonlinear Energy Sink
by Xiang Fu, Sha Wei, Hu Ding and Li-Qun Chen
Appl. Sci. 2025, 15(3), 1380; https://doi.org/10.3390/app15031380 - 29 Jan 2025
Abstract
Reductions in the vibration of a continuum system via a nonlinear energy sink have been widely investigated. It is usually assumed that weight effects can be ignored if the vibration is measured from the static equilibrium configuration. The present investigation reveals the dynamic [...] Read more.
Reductions in the vibration of a continuum system via a nonlinear energy sink have been widely investigated. It is usually assumed that weight effects can be ignored if the vibration is measured from the static equilibrium configuration. The present investigation reveals the dynamic effects of weight on the vertical transverse vibrations of a Euler–Bernoulli beam coupled with a nonlinear energy sink. The governing equations considering and neglecting weights were derived. The equations were discretized with some numerical support. The discretized equations were analytically solved via the harmonic balance method. The harmonic balance solutions were compared with the numerical solution via the Runge–Kutta method. Finite element simulations were performed via ANSYS software (version number: 2.2.1). Free and forced vibrations, predicted by equations considering or neglecting the weights, were compared with the finite element solutions. For the forced vibrations, the amplitude–frequency responses determined by the harmonic balance method agree well with those calculated by the Runge–Kutta method. The free and forced vibration responses predicted by the equations considering the weights are closer to those computed by the finite element method than the responses predicted by the equation neglecting the weights. The assumption that weights can be balanced by static deflections leads to errors in the analysis of the vertical transverse vibrations of a Euler–Bernoulli beam with a nonlinear energy sink. Full article
(This article belongs to the Special Issue Advances in Architectural Acoustics and Vibration)
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<p>A mechanical model of an elastic beam coupled with a nonlinear energy sink.</p>
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<p>Time history curve of free vibration when nonlinear energy sink weight is considered, with different truncated orders: (<b>a</b>) displacement of 1/8 point when nonlinear energy sink is placed at 1/2 point of beam; (<b>b</b>) displacement of 1/4 point when nonlinear energy sink is placed at 1/2 point of beam.</p>
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<p>Time history curve of free vibration when nonlinear energy sink weight is neglected, with different truncated orders: (<b>a</b>) displacement of 1/8 point when nonlinear energy sink is placed at 1/2 point of beam; (<b>b</b>) displacement of 1/4 point when nonlinear energy sink is placed at 1/2 point of beam.</p>
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<p>Amplitude−frequency response of forced vibration when nonlinear energy sink weight is considered, with different orders of Galerkin truncation: (<b>a</b>) 1/8 point of beam when nonlinear energy sink is placed at 1/2 point of beam, (<b>b</b>) 1/4 point of beam when nonlinear energy sink is placed at 1/2 point of beam.</p>
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<p>Amplitude−frequency response of forced vibration when nonlinear energy sink weight is neglected, with different orders of Galerkin truncation: (<b>a</b>) 1/8 point of beam when nonlinear energy sink is placed at 1/2 point of beam, (<b>b</b>) 1/4 point of beam when nonlinear energy sink is placed at 1/2 point of beam.</p>
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<p>Comparison of amplitude−frequency curves of analytic solution and numerical solution when nonlinear energy sink weight is considered: (<b>a</b>) 1/8 point of beam when nonlinear energy sink is placed at 1/2 point of beam, (<b>b</b>) 1/4 point of beam when nonlinear energy sink is placed at 1/2 point of beam.</p>
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<p>Comparison of amplitude−frequency curves of analytic solution and numerical solution when nonlinear energy sink weight is neglected: (<b>a</b>) 1/8 point of beam when nonlinear energy sink is placed at 1/2 point of beam, (<b>b</b>) 1/4 point of beam when nonlinear energy sink is placed at 1/2 point of beam.</p>
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<p>Time history curve of free vibration: (<b>a</b>) response of 1/8 point of beam when nonlinear energy sink is placed at 1/2 point of beam; (<b>b</b>) response of 1/4 point of beam when nonlinear energy sink is placed at 1/2 point of beam.</p>
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<p>Amplitude Spectrum of free vibration: (<b>a</b>) response of 1/8 point of beam when nonlinear energy sink is placed at 1/2 point of beam; (<b>b</b>) response of 1/4 point of beam when nonlinear energy sink is placed at 1/2 point of beam.</p>
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<p>Time history curve of forced vibration with first−order excitation frequency: (<b>a</b>) response of 1/8 point of beam when nonlinear energy sink is placed at 1/2 point of beam; (<b>b</b>) response of 1/4 point of beam when nonlinear energy sink is placed at 1/2 point of beam.</p>
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<p>Amplitude−frequency response of forced vibration with first−order excitation frequency: (<b>a</b>) response of 1/8 point of beam when nonlinear energy sink is placed at 1/2 point of beam; (<b>b</b>) response of 1/4 point of beam when nonlinear energy sink is placed at 1/2 point of beam.</p>
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<p>Time history curve of forced vibration with second−order excitation frequency: (<b>a</b>) response of 1/8 point of beam when the nonlinear energy sink is placed at 1/2 point of beam; (<b>b</b>) response of 1/4 point of beam when nonlinear energy sink is placed at 1/2 point of beam.</p>
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<p>Amplitude−frequency response of forced vibration with second−order excitation frequency: (<b>a</b>) response of 1/8 point of beam when nonlinear energy sink is placed at 1/2 point of beam; (<b>b</b>) response of 1/4 point of beam when nonlinear energy sink is placed at 1/2 point of beam.</p>
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<p>Amplitude spectra of free vibration of ANSYS software (version number:2.2.1) simulation and Runge−Kutta method when nonlinear energy sink is placed at 1/2 point of beam: (<b>a</b>) frequency response of 1/8 point of beam; (<b>b</b>) frequency response of 1/4 point of beam.</p>
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<p>Results of ANSYS software (version number:2.2.1) simulation and Runge−Kutta method when nonlinear energy sink is placed at 1/2 point of beam: (<b>a</b>) frequency response of 1/8 point of beam; (<b>b</b>) frequency response of 1/4 point of beam.</p>
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<p>Results of ANSYS software (version number:2.2.1) simulation and harmonic balance method when nonlinear energy sink is placed at 1/2 point of beam: (<b>a</b>) frequency response of 1/8 point of beam; (<b>b</b>) frequency response of 1/4 point of beam.</p>
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<p>Results of ANSYS software (version number:2.2.1) simulation and Runge−Kutta method when nonlinear energy sink is placed at 1/2 point of beam with second−order primary frequency excitation: (<b>a</b>) frequency response of 1/8 point of beam; (<b>b</b>) frequency response of 1/4 point of beam.</p>
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<p>Results of ANSYS software (version number:2.2.1) simulation and harmonic balance method when nonlinear energy sink is placed at 1/2 point of beam with second-order primary frequency excitation: (a) frequency response of 1/8 point of beam; (b) frequency response of 1/4 point of beam.</p>
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25 pages, 3346 KiB  
Article
Study on Dynamic Response of Damping Type Composite Floor Slabs Considering Interlayer Interaction Influences
by Liangming Sun, Ting Xu, Feng Tian, Yijie Zhang, Hanbing Zhao and Aziz Hasan Mahmood
J. Compos. Sci. 2025, 9(2), 57; https://doi.org/10.3390/jcs9020057 - 26 Jan 2025
Abstract
In order to explore the vibration mechanism of vibration damping composite floor slabs and further enrich the theory of floor slab vibration calculation, the free vibration characteristics of vibration damped composite floor slabs and the dynamic response of vibration damped composite floor slabs [...] Read more.
In order to explore the vibration mechanism of vibration damping composite floor slabs and further enrich the theory of floor slab vibration calculation, the free vibration characteristics of vibration damped composite floor slabs and the dynamic response of vibration damped composite floor slabs under multi-source excitation is analyzed using first type Chebyshev polynomials to construct the displacement function and derive an analytical solution. The three-dimensional laminated theory is employed, considering the interlayer interaction. Based on the proposed method, the influences of loading types, positions, magnitudes, and frequencies on the vertical vibration of floor slabs are calculated. The study illustrates that, under the action of multi-source excitation, the displacement and acceleration responses calculated by the method proposed in this paper are always greater than those calculated by the single-plate theoretical solution. The dynamic responses of the vibration damping composite floor slab decrease with the increase of the thickness and elastic modulus of the vibration damping layer. Under different thicknesses of the vibration damping layer, the peak accelerations of the vibration damping composite floor slabs increase linearly with the growth of the load amplitude. In addition, the load movement path has a significant effect on the vibration response of the floor slab. When the moving load moves along the short side of the floor, the displacement response of the floor is generally greater than that along the long side of the floor. Full article
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<p>Analysis model and coordinate system of vibration damped composite floor slab.</p>
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<p>Modal vibration modes of the composite floor slab with four-sides simply supported.</p>
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<p>Layout diagrams of fixed loads and moving loads.</p>
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<p>Comparison of the vertical vibration responses at the center point of the floor slab under harmonic loads. (<b>a</b>) displacement, (<b>b</b>) acceleration.</p>
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<p>Comparison of the vertical vibration responses at the center point of the floor slab under moving loads.</p>
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<p>Comparison of vertical vibration response at the center point of a vibration damped composite floor slab under multi-source excitation.</p>
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<p>Effect of damping layer thickness on dynamic response of composite floor slabs.</p>
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<p>Effect of elastic modulus of damping layer on the dynamic response of composite floor slabs.</p>
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<p>The influence of harmonic loads on the dynamic response of composite floor slabs.</p>
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<p>The influence of the amplitude of moving loads on the dynamic response of composite floor slabs in buildings.</p>
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<p>The influence of the load velocity on the peak acceleration responses of composite floor slabs.</p>
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<p>Layout diagram of various loads.</p>
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<p>Floor displacement response under different load positions.</p>
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16 pages, 3334 KiB  
Article
Lead-Free Ceramics in Prestressed Ultrasonic Transducers
by Claus Scheidemann, Peter Bornmann, Walter Littmann and Tobias Hemsel
Actuators 2025, 14(2), 55; https://doi.org/10.3390/act14020055 - 25 Jan 2025
Viewed by 190
Abstract
Today’s ultrasonic transducers find broad application in diverse technology branches and most often cannot be replaced by other actuators. They are typically based on lead-containing piezoelectric ceramics. These should be replaced for environmental and health issues by lead-free alternatives. Multiple material alternatives are [...] Read more.
Today’s ultrasonic transducers find broad application in diverse technology branches and most often cannot be replaced by other actuators. They are typically based on lead-containing piezoelectric ceramics. These should be replaced for environmental and health issues by lead-free alternatives. Multiple material alternatives are already known, but there is a lack of information about their technological readiness level. To fill this gap, a small series of prestressed longitudinally vibrating transducers was set up with a standard PZT material and two lead-free variants within this study. The entire process for building the transducers is documented: characteristics of individual ring ceramics, burn-in results, and free vibration and characteristics under load are shown. The main result is that the investigated lead-free materials are ready to use within ultrasonic bolted Langevin transducers (BLTs) for medium-power applications, when the geometrical setup of the transducer is adopted. Since lead-free ceramics need higher voltages to achieve the same power level, the driving electronics or the mechanical setup must be altered specifically for each material. Lower self-heating of the lead-free materials might be attractive for heat-sensitive processes. Full article
(This article belongs to the Special Issue Piezoelectric Ultrasonic Actuators and Motors)
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<p>BLT design: The backing mass (③), 4 piezoelectric ceramic rings (④), and 5 electrodes (②) are prestressed by a hollow screw (①); the base body (⑤) has an internal thread at its front side to attach a sonotrode (not used within this study).</p>
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<p>Experimental setup: Structure of the actuator in a loaded state (immersed in water) and its electrical control, with a schematic representation of the water circuit and non-contact measurement of the surface temperature of the piezo-elements. Dashed half-circles below the transducer’s front-face symbolize acoustic waves being radiated into water. The dashed lines inside the transducer indicate the central channel for liquid throughput.</p>
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<p>Series of assembled transducers with PZT and lead-free piezoelectric ceramics, plastic mounting bracket, and ATHENA ultrasound generator being used for resonance-controlled operation and measurements.</p>
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<p>Measured frequency responses of the short-circuit input admittance of 50 individual ring ceramics (coloured solid lines) compared to simulation results (dashed black line) for (<b>a</b>) PIC 181, (<b>b</b>) PIC 758, and (<b>c</b>) PIC HQ2. Please refer to the different frequency ranges used in the diagrams for convenience.</p>
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<p>Results of burn-in process (free resonant vibration at ≈1 m/s for ≈6 min): (<b>a</b>) Temperature rise over time, (<b>b</b>) resonance frequency change over temperature rise, (<b>c</b>) voltage related to tip velocity over time, (<b>d</b>) motional current related to tip velocity over temperature rise.</p>
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<p>Small-signal admittance characteristics (free vibration, room temperature).</p>
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<p>Results of tests (resonance-controlled continuous vibration up to 1.5 m/s, heat up over time): (<b>a</b>) Dependency of tip velocity and motional current, (<b>b</b>) mechanical damping factor over tip velocity, (<b>c</b>) ratio of motional current and tip velocity over temperature rise, and (<b>d</b>) mechanical damping factor over temperature rise.</p>
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<p>Results of short-time operation tests; colours stand for different materials (PIC 181: red, PIC 758: blue, PIC HQ2: green): (<b>a</b>) Dependency of tip velocity and motional current for the no-load vibration of the transducer being fixed in the clamping, and (<b>b</b>) active power over motional current at free vibration (dark colours) and under water load (light colours).</p>
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<p>Results of load tests with heating up (immersion of transducer tip into water, controlled vibration at different levels of motional current, continuous drive until steady state temperature): (<b>a</b>) motional current amplitude over time, (<b>b</b>) temperature rise over time, (<b>c</b>) voltage amplitude over time, and (<b>d</b>) active power over motional current amplitude.</p>
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14 pages, 3007 KiB  
Article
The Potential of a Thermoelectric Heat Dissipation System: An Analytical Study
by Xuechun Li, Rujie Shi and Kang Zhu
Energies 2025, 18(3), 555; https://doi.org/10.3390/en18030555 - 24 Jan 2025
Viewed by 291
Abstract
Thermoelectric heat dissipation systems offer unique advantages over conventional systems, including vibration-free operation, environmental sustainability, and enhanced controllability. This study examined the benefits of incorporating a thermoelectric cooler (TEC) into conventional heat sinks and investigated strategies to improve heat dissipation efficiency. A theoretical [...] Read more.
Thermoelectric heat dissipation systems offer unique advantages over conventional systems, including vibration-free operation, environmental sustainability, and enhanced controllability. This study examined the benefits of incorporating a thermoelectric cooler (TEC) into conventional heat sinks and investigated strategies to improve heat dissipation efficiency. A theoretical model introducing a dimensionless evaluation index (rq) is proposed to assess the system’s performance, which measures the ratio of the heat dissipation density of a conventional heat dissipation system to that of a thermoelectric heat dissipation system. Here, we subjectively consider 0.9 as a cutoff, and when rq<0.9, the thermoelectric heat dissipation system shows substantial superiority over conventional ones. In contrast, for rq>0.9, the advantage of the thermoelectric system weakens, making conventional systems more attractive. This analysis examined the effects of engineering leg length (L*), the heat transfer allocation ratio (rh), and temperature difference (ΔT) on heat dissipation capabilities. The results indicated that under a fixed heat source temperature, heat sink temperature, and external heat transfer coefficient, an optimal engineering leg length exists, maximizing the system’s heat dissipation performance. Furthermore, a detailed analysis revealed that the thermoelectric system demonstrated exceptional performance under small temperature differences, specifically when the temperature difference was below 32 K with the current thermoelectric (TE) materials. For moderate temperature differences between 32 K and 60 K, the system achieved optimal performance when rh2.4+1.37e0.019ΔT. This work establishes a theoretical foundation for applying thermoelectric heat dissipation systems and provides valuable insights into optimizing hybrid heat dissipation systems. Full article
(This article belongs to the Special Issue Recent Advances in Thermoelectric Energy Conversion)
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<p>(<b>A</b>) The conventional heat dissipation system model. (<b>B</b>) The thermal network and qualitative temperature allocation of the conventional heat dissipation system.</p>
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<p>(<b>A</b>) The thermoelectric heat dissipation system model. (<b>B</b>) Thermal network and qualitative temperature allocation of thermoelectric heat dissipation system.</p>
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<p>Comparison between the theoretical model calculation results of this study and the literature results under two different thermal resistance conditions: (<b>A</b>) <span class="html-italic">R<sub>t,h</sub></span>1 = 0.03823 K/W and (<b>B</b>) <span class="html-italic">R<sub>t,h</sub></span>2 = 0.15186 K/W.</p>
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<p>(<b>A</b>) Validations of correlations of <span class="html-italic">q<sub>h,max</sub></span>. (<b>B</b>) The optimal engineering leg length under different total heat transfer coefficients. (<b>C</b>) The optimal engineering leg length under different <span class="html-italic">r<sub>h</sub></span> conditions when <span class="html-italic">h<sub>h</sub></span> = 1000 W/m<sup>2</sup>·K. (<b>D</b>) The optimal engineering leg length under different <span class="html-italic">r<sub>h</sub></span> conditions when <span class="html-italic">h<sub>h</sub></span> = 10,000 W/m<sup>2</sup>·K.</p>
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<p>The contour of influencing parameters (<b>A</b>) <math display="inline"><semantics> <mrow> <msub> <mi>β</mi> <mn>1</mn> </msub> </mrow> </semantics></math>; (<b>B</b>) <math display="inline"><semantics> <mrow> <msub> <mi>β</mi> <mn>2</mn> </msub> </mrow> </semantics></math>; (<b>C</b>) K; (<b>D</b>) <span class="html-italic">r<sub>q</sub></span>.</p>
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<p>The distribution curve of heat dissipation density for both with TE and without TE.</p>
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<p>(<b>A</b>) Dimensionless evaluation index (<span class="html-italic">r<sub>q</sub></span>) under different <span class="html-italic">r<sub>h</sub> </span>and ΔT. (<b>B</b>) The relationship between <span class="html-italic">r<sub>h</sub></span> and <math display="inline"><semantics> <mrow> <mi mathvariant="sans-serif">Δ</mi> <mi>T</mi> </mrow> </semantics></math> under <span class="html-italic">r<sub>q</sub></span> = 0.9.</p>
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29 pages, 3225 KiB  
Article
Hierarchical Free Vibration Analysis of Variable-Angle Tow Shells Using Unified Formulation
by Domenico Andrea Iannotta, Gaetano Giunta, Levent Kirkayak and Marco Montemurro
J. Compos. Sci. 2025, 9(2), 55; https://doi.org/10.3390/jcs9020055 - 24 Jan 2025
Viewed by 248
Abstract
This paper investigates the dynamic behavior of shell structures presenting variable-angle tow laminations. The choice of placing fibers along curvilinear patterns allows for a broader structural design space, which is advantageous in several engineering contexts, provided that more complex numerical analyses are managed. [...] Read more.
This paper investigates the dynamic behavior of shell structures presenting variable-angle tow laminations. The choice of placing fibers along curvilinear patterns allows for a broader structural design space, which is advantageous in several engineering contexts, provided that more complex numerical analyses are managed. In this regard, Carrera’s unified formulation has been widely used for studying variable-angle tow plates and shells. This article aims to expand this formulation through the derivation of the complete formulation for a generic shell reference surface. The principle of virtual displacements is used as a variational statement for obtaining, in a weak sense, the stiffness and mass matrices within the finite element solution method. The free vibration problem of singly and doubly curved variable-angle tow shells is then addressed. The proposed approach is compared to Abaqus three-dimensional reference solutions and classical theories to investigate the effectiveness of the developed models in predicting the vibrational frequencies and modes. The results demonstrate a good agreement between the proposed approach and reference solutions. Full article
(This article belongs to the Special Issue Feature Papers in Journal of Composites Science in 2024)
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<p>Shell geometry and reference system.</p>
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<p>Example of in-plane fibers’ path.</p>
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<p>Acronym system.</p>
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<p>Tying points for the CUF MITC9 shell element.</p>
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<p>Geometry and reference system, case 1.</p>
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<p>Fibers’ orientation, case 1.</p>
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<p>Fibers’ orientation, case 2.</p>
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<p>First six eigenmodes, <math display="inline"><semantics> <mrow> <msub> <mi>l</mi> <mi>α</mi> </msub> <mo>/</mo> <mi>h</mi> <mo>=</mo> <mn>10</mn> </mrow> </semantics></math>, case 2.</p>
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<p>Geometry and reference system, case 3.</p>
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<p>Stacking sequence, case 3.</p>
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<p>First six eigenmodes, <math display="inline"><semantics> <mrow> <msub> <mi>l</mi> <mi>α</mi> </msub> <mo>/</mo> <mi>h</mi> <mo>=</mo> <mn>10</mn> </mrow> </semantics></math>, case 3.</p>
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19 pages, 3413 KiB  
Article
Free Vibrations and Flutter Analysis of Composite Plates Reinforced with Carbon Nanotubes
by Małgorzata Chwał
Appl. Sci. 2025, 15(3), 1140; https://doi.org/10.3390/app15031140 - 23 Jan 2025
Viewed by 310
Abstract
This paper considers the free vibration and flutter of carbon nanotube (CNT) reinforced nanocomposite plates subjected to supersonic flow. From the literature review, a great deal of research has been conducted on the free vibration and flutter response of high-volume CNT/nanocomposite structures; however, [...] Read more.
This paper considers the free vibration and flutter of carbon nanotube (CNT) reinforced nanocomposite plates subjected to supersonic flow. From the literature review, a great deal of research has been conducted on the free vibration and flutter response of high-volume CNT/nanocomposite structures; however, there is little research on the flutter instability of low-volume CNT/nanocomposite structures. In this study, free vibration and flutter analysis of classical CNT/nanocomposite thin plates with aligned and uniformly distributed reinforcement and low CNT volume fraction are performed. The geometry of the CNTs and the definition of the nanocomposite material properties are considered. The nanocomposite properties are estimated based on micromechanical modeling, while the governing relations of the nanocomposite plates are derived according to Kirchhoff’s plate theory with von Karman nonlinear strains. Identification of vibrational modes for nanocomposite thin plates and analytical/graphical evaluation of flutter are presented. The novel contribution of this work is the analysis of the eigenfrequencies and dynamic instabilities of nanocomposite plates with a low fraction of CNTs aligned and uniformly distributed in the polymer matrix. This article is helpful for a comprehensive understanding of the influence of a low-volume fraction and uniform distribution of CNTs and boundary conditions on the dynamic instabilities of nanocomposite plates. Full article
(This article belongs to the Section Acoustics and Vibrations)
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<p>The concept of the micromechanical modeling of the equivalent nanocomposite with aligned carbon nanotubes.</p>
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<p>The mechanical model of the equivalent nanocomposite plate subjected to a supersonic airflow in the <span class="html-italic">x</span>-direction. The <span class="html-italic">z</span>-coordinate is determined from the mid-plane of the plate.</p>
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<p>Variation of the non-dimensional fundamental frequency of the simply supported nanocomposite plate as a function of the aspect ratio of the plates (<span class="html-italic">h</span> = 5 mm, <span class="html-italic">m</span> = <span class="html-italic">n</span> = 1) for various volume fractions of CNTs (1%, 3%, and 5%).</p>
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<p>Variation of the non-dimensional fundamental frequency of the simply supported nanocomposite plate as a function of the volume fractions of CNTs for the various aspect ratios of the plates (<span class="html-italic">h</span> = 5 mm, <span class="html-italic">m</span> = <span class="html-italic">n</span> = 1).</p>
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<p>Variation of the fundamental frequency of simply supported nanocomposite plate versus the plate aspect ratio for various plate thicknesses (<span class="html-italic">h</span> = 2 mm, and <span class="html-italic">h</span> = 5 mm) and volume fractions of CNTs (1%, 3%, and 5%).</p>
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<p>Variation of the non-dimensional frequency of simply supported square nanocomposite plates (<span class="html-italic">a/b</span> =1, <span class="html-italic">a</span>/<span class="html-italic">h</span> = 200) versus the volume fraction of CNTs for the first five modes.</p>
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<p>Variation of the non-dimensional frequency of simply supported rectangular nanocomposite plates (<span class="html-italic">a/b</span> = 2, <span class="html-italic">a</span>/<span class="html-italic">h</span> = 200) versus the volume fraction of CNTs for the first five modes.</p>
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<p>The eigenvalue curves for 1% vol. CNT/nanocomposite plate for simply supported case (upper curve) and clamped case (lower curve) for <span class="html-italic">a/b</span> = 1, and <span class="html-italic">h</span> = 2 mm.</p>
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<p>The eigenvalue curves for isotropic plate (<span class="html-italic">k</span> = 1) and 1% CNT/nanocomposite orthotropic plate (<span class="html-italic">k</span> = 0.257) for <span class="html-italic">a/b</span> = 1, and <span class="html-italic">h</span> = 2 mm.</p>
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<p>The eigenvalue curves for various volume fractions of CNTs in nanocomposite plates–SSSS, <span class="html-italic">a/b</span> = 1, <span class="html-italic">h</span> = 2 mm.</p>
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<p>Non-dimensional pressure <span class="html-italic">β</span> (dashed line) and non-dimensional frequency <span class="html-italic">λ</span> (solid line) for 1% CNT/nanocomposite plate– <span class="html-italic">a/b</span> = 1, <span class="html-italic">a</span>/<span class="html-italic">h</span> = 500, <span class="html-italic">h</span> = 2 mm.</p>
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<p>Graphical construction of the flutter problem solution—the coalescence point for 1% CNT/nanocomposite–SSSS square plate.</p>
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20 pages, 1816 KiB  
Article
Accurate Cardiac Duration Detection for Remote Blood Pressure Estimation Using mm-Wave Doppler Radar
by Shengze Wang, Mondher Bouazizi, Siyuan Yang and Tomoaki Ohtsuki
Sensors 2025, 25(3), 619; https://doi.org/10.3390/s25030619 - 21 Jan 2025
Viewed by 437
Abstract
This study introduces a radar-based model for estimating blood pressure (BP) in a touch-free manner. The model accurately detects cardiac activity, allowing for contactless and continuous BP monitoring. Cardiac motions are considered crucial components for estimating blood pressure. Unfortunately, because these movements are [...] Read more.
This study introduces a radar-based model for estimating blood pressure (BP) in a touch-free manner. The model accurately detects cardiac activity, allowing for contactless and continuous BP monitoring. Cardiac motions are considered crucial components for estimating blood pressure. Unfortunately, because these movements are extremely subtle and can be readily obscured by breathing and background noise, accurately detecting these motions with a radar system remains challenging. Our approach to radar-based blood pressure monitoring in this research primarily focuses on cardiac feature extraction. Initially, an integrated-spectrum waveform is implemented. The method is derived from the short-time Fourier transform (STFT) and has the ability to capture and maintain minute cardiac activities. The integrated spectrum concentrates on energy changes brought about by short and high-frequency vibrations, in contrast to the pulse-wave signals used in previous works. Hence, the interference caused by respiration, random noise, and heart contractile activity can be effectively eliminated. Additionally, we present two approaches for estimating cardiac characteristics. These methods involve the application of a hidden semi-Markov model (HSMM) and a U-net model to extract features from the integrated spectrum. In our approach, the accuracy of extracted cardiac features is highlighted by the notable decreases in the root mean square error (RMSE) for the estimated interbeat intervals (IBIs), systolic time, and diastolic time, which were reduced by 87.5%, 88.7%, and 73.1%. We reached a comparable prediction accuracy even while our subject was breathing normally, despite previous studies requiring the subject to hold their breath. The diastolic BP (DBP) error of our model is 3.98±5.81 mmHg (mean absolute difference ± standard deviation), and the systolic BP (SBP) error is 6.52±7.51 mmHg. Full article
(This article belongs to the Special Issue Analyzation of Sensor Data with the Aid of Deep Learning)
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<p>An illustration of (<b>a</b>) conventional assumption on systolic and diastolic timing extraction and (<b>b</b>) actual systolic and diastolic timings from ECG waveform.</p>
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<p>The system model and the setup of the Doppler sensor for capturing the heartbeat signal.</p>
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<p>An illustration of the heart parts.</p>
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<p>Flowchart of the proposed method.</p>
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<p>(<b>a</b>) The spectrogram of the conventional pulse wave signal; (<b>b</b>) the spectrogram of the higher-frequency radar signal selected in our work.</p>
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<p>Integrated spectrum and pulse wave with ECG as gold standard.</p>
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<p>Example of an HSMM algorithm.</p>
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<p>The estimated state change generated by HSMM.</p>
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<p>Example of a U-net structure.</p>
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<p>An illustration of (<b>A</b>) shows the label are generated from the ECG signal. (<b>B</b>) shows the comparison between Integrated spectrum and the generated label.</p>
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<p>The integrated spectrum, state changes, and U-net results compared with the R-peaks and end of the T-wave in the ECG.</p>
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