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Keywords = wire bonding

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22 pages, 6412 KiB  
Article
Numerical Analysis and Theoretical Study on the Interfacial Bonding Behavior of High-Strength Steel Stainless Wire Mesh-Reinforced ECC and Concrete
by Chao Li, Yao Zou, Ziyuan Li, Xuyan Zou, Ke Li, Juntao Zhu, Hongbo Xiao and Jianwei Fan
Materials 2024, 17(23), 5912; https://doi.org/10.3390/ma17235912 - 3 Dec 2024
Viewed by 389
Abstract
In order to investigate the interfacial bonding properties of high-strength steel stainless wire mesh-reinforced ECC (HSSWM-ECC) and concrete, a finite element model was established for two types of interfaces based on experimental research. The results show that the failure modes observed in the [...] Read more.
In order to investigate the interfacial bonding properties of high-strength steel stainless wire mesh-reinforced ECC (HSSWM-ECC) and concrete, a finite element model was established for two types of interfaces based on experimental research. The results show that the failure modes observed in the 21 groups of simulations can be classified into three categories: debonding failure, ECC extrusion failure and concrete splitting failure. The failure mode was mainly affected by the type of interface. The effective anchorage length is inversely proportional to the strength of the concrete and proportional to the stiffness and thickness of the HSSWM-ECC. The capacity of the roughening interface is positively correlated with the concrete strength and bonding length, but negatively correlated with the interfacial width ratio. Increasing both the number and width of grooves within the effective range enhances the interfacial capacity, whereas higher concrete strengths tend to reduce it. Based on the above results, calculation models for the effective anchorage length and bearing capacity were established separately for the two types of interfaces. The theoretical model for the interfacial bonding property between HSSWM-ECC and concrete has been refined. These advancements establish a theoretical groundwork for the design of concrete structures strengthened with HSSWM-ECC. Full article
(This article belongs to the Section Materials Simulation and Design)
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Figure 1
<p>Geometry of specimen (Unit: mm).</p>
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<p>Processing modes of interface (<b>a</b>) Manual roughening; (<b>b</b>) High-pressure water flushing; (<b>c</b>) Mechanical grooving.</p>
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<p>Loading device.</p>
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<p>Interfacial load–slip curves of main specimens.</p>
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<p>Constitutive models of concrete: (<b>a</b>) Tensile constitutive model of concrete; (<b>b</b>) Compressive constitutive model of concrete.</p>
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<p>Constitutive models of ECC: (<b>a</b>) Tensile constitutive model of ECC; (<b>b</b>) Compressive constitutive model of ECC.</p>
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<p>Constitutive model of steel wire.</p>
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<p>Mechanical model of bonding interfaces: (<b>a</b>) Interface I; (<b>b</b>) Interface II.</p>
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<p>Model of Interface I: (<b>a</b>) Zero-thickness interface unit; (<b>b</b>) Local bond-slip model of the HSSWM-ECC and concrete.</p>
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<p>Model of Interface II.</p>
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<p>FE model of specimen: (<b>a</b>) Geometric view 1; (<b>b</b>) Geometric view 2; (<b>c</b>) meshing.</p>
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<p>Comparison of experimental and simulated F–S curves (<b>a</b>) A2; (<b>b</b>) C1; (<b>c</b>) C2; (<b>d</b>) D2.</p>
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<p>Characteristics of groove (unit: mm).</p>
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<p>Interfacial stress distribution under peak load: (<b>a</b>) Debonding failure of interface; (<b>b</b>) ECC extrusion failure; (<b>c</b>) Concrete splitting failure.</p>
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<p>Load–slip curves of specimens with Interface I: (<b>a</b>) Group Z; (<b>b</b>) Group Y; (<b>c</b>) Group X.</p>
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<p>Load–slip curves of specimens with Interface II: (<b>a</b>) Group W; (<b>b</b>) Group V; (<b>c</b>) Group U.</p>
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<p>Model of interfacial load–slip curve: (<b>a</b>) Typical curve; (<b>b</b>) Model of load–slip relationship.</p>
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<p>Interfacial stress distribution curves of specimens with Interface I under each load: (<b>a</b>) Z1; (<b>b</b>) Z2; (<b>c</b>) Z3; (<b>d</b>) Z4; (<b>e</b>) Y1; (<b>f</b>) Y2; (<b>g</b>) Y3; (<b>h</b>) Y4; (<b>i</b>) X1; (<b>j</b>) X2; (<b>k</b>) X3.</p>
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<p>Interfacial stress distribution curves of specimens with Interface II under each load: (<b>a</b>) W1; (<b>b</b>) W2; (<b>c</b>) W3; (<b>d</b>) W4; (<b>e</b>) V1; (<b>f</b>) V2; (<b>g</b>) V3; (<b>h</b>) U1; (<b>i</b>) U2; (<b>j</b>) U3.</p>
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<p>Acquisition of effective anchorage length.</p>
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<p>Comprehensive adjustment coefficient KH linear regression analysis.</p>
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<p>Comprehensive adjustment coefficient <span class="html-italic">k</span><sub>G</sub> linear regression analysis.</p>
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19 pages, 14262 KiB  
Article
Study on the Microstructure and Properties of Al Alloy/Steel CMT Welding–Brazing Joints Under Different Pulse Magnetic Field Intensities
by Juan Pu, Tingmu Chen, Jiawei Rao, Yubo Sun, Yunxia Chen, Changhua Jiang and Huawei Sun
Coatings 2024, 14(12), 1515; https://doi.org/10.3390/coatings14121515 - 30 Nov 2024
Viewed by 509
Abstract
Butt welding experiments on 6061 Al alloy and Q235B steel of 2 mm thickness were conducted using an ER4047F flux-cored wire as the filler metal, after adding a pulsed magnetic field into the process of cold metal transfer (CMT) welding. The effect of [...] Read more.
Butt welding experiments on 6061 Al alloy and Q235B steel of 2 mm thickness were conducted using an ER4047F flux-cored wire as the filler metal, after adding a pulsed magnetic field into the process of cold metal transfer (CMT) welding. The effect of the pulsed magnetic field intensity on the macro morphology, microstructure, tensile strength and corrosion resistance of the welding–brazing joint was analyzed. The results showed that when the pulsed magnetic field intensity increased from 0 to 60 mT, the wettability and spreadability of the liquid metal were improved. As a result, the appearance of the Al alloy/steel joint was nice. However, when the pulsed magnetic field intensity was 80 mT, the stability of the arc and the forming quality of the joint decreased, which resulted in a deterioration in the appearance of the joint. A pulsed magnetic field with different intensities did not alter the microstructure of the joint. All of the joint was composed of θ-Fe2(Al,Si)5 and τ5-Al7.2Fe1.8Si at the interface and Al-Si eutectic phase and α-Al solid solution at the weld seam zone. Actually, with the pulsed magnetic field intensity increasing from 0 mT to 60 mT, the IMC thickness in the interfacial layer gradually reduced under the action of electromagnetic stirring. Also, the grain in the weld seam was refined, and elements were distributed uniformly. But when the pulsed magnetic field intensity was 80 mT, the grain in the weld seam began to coarsen, and the intermetallic compound (IMC) thickness was too small, which was unfavorable for the metallurgical bonding of Al alloy and steel. Therefore, with the increase in pulsed magnetic field intensity, the tensile strength of the joints first increased and then decreased, and it reached its maximum of 187.7 MPa with a pulsed magnetic field intensity of 60 mT. Similarly, the corrosion resistance of the joint first increased and then decreased, and it was best when the pulse magnetic field intensity was 60 mT. The Nyquist plot and Bode plot confirmed this result. The addition of a pulsed magnetic field caused less fluctuation in the anode current density, resulting in less localized corrosion of the joint using the scanning vibrating electrode technique (SVET). The XPS analysis showed the Al-Fe-Si compounds replacing the Fe-Al compounds in the joint was the main reason for improving its corrosion resistance under the action of a pulsed magnetic field. Full article
(This article belongs to the Special Issue Laser Surface Engineering and Additive Manufacturing)
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Figure 1
<p>Schematic diagram of pulsed magnetic field generator-assisted CMT welding–brazing: (<b>a</b>) schematic diagram; (<b>b</b>) working diagram of electromagnet in pulse period (4T/5); (<b>c</b>) working diagram of electromagnet in pulse period (T).</p>
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<p>Surface appearance and cross-section morphology of joints under different magnetic field intensities: (<b>a<sub>1</sub></b>–<b>a<sub>3</sub></b>) 0 mT; (<b>b<sub>1</sub></b>–<b>b<sub>3</sub></b>) 20 mT; (<b>c<sub>1</sub></b>–<b>c<sub>3</sub></b>) 40 mT; (<b>d<sub>1</sub></b>–<b>d<sub>3</sub></b>) 60 mT; (<b>e<sub>1</sub></b>–<b>e<sub>3</sub></b>) 80 mT.</p>
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<p>Microstructural characterization of joints under different magnetic field intensities: (<b>a<sub>1</sub></b>–<b>a<sub>3</sub></b>) 0 mT; (<b>b<sub>1</sub></b>–<b>b<sub>3</sub></b>) 20 mT; (<b>c<sub>1</sub></b>–<b>c<sub>3</sub></b>) 40 mT; (<b>d<sub>1</sub></b>–<b>d<sub>3</sub></b>) 60 mT; (<b>e<sub>1</sub></b>–<b>e<sub>3</sub></b>) 80 mT.</p>
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<p>SEM and EDS line scanning results of IMC layer at the steel side under different magnetic field intensities: (<b>a<sub>1</sub></b>,<b>a<sub>2</sub></b>) 20 mT; (<b>b<sub>1</sub></b>,<b>b<sub>2</sub></b>) 40 mT; (<b>c<sub>1</sub></b>,<b>c<sub>2</sub></b>) 60 mT; (<b>d<sub>1</sub></b>,<b>d<sub>2</sub></b>) 80 mT.</p>
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<p>SEM area scanning results of joints under different magnetic field intensities: (<b>a</b>) 20 mT; (<b>b</b>) 40 mT; (<b>c</b>) 60 mT; (<b>d</b>) 80 mT.</p>
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<p>XRD results of joints under different magnetic field intensities: (<b>a</b>) without magnetic field; (<b>b</b>) 60 mT.</p>
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<p>The effect of pulsed magnetic field intensity on the tensile strength of joints.</p>
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<p>Fracture location and fracture morphology of the joint under different magnetic field intensities: (<b>a<sub>1</sub></b>,<b>a<sub>2</sub></b>) 20 mT; (<b>b<sub>1</sub></b>,<b>b<sub>2</sub></b>) 40 mT; (<b>c<sub>1</sub></b>,<b>c<sub>2</sub></b>) 60 mT; (<b>d<sub>1</sub></b>,<b>d<sub>2</sub></b>) 80 mT.</p>
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<p>Polarization curves of joints under different pulsed magnetic field intensities.</p>
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<p>Electrochemical impedance spectroscopy (EIS) spectra of joints under different magnetic field intensities: (<b>a</b>) Nyquist diagram; (<b>b</b>) mode frequency Bode diagram; (<b>c</b>) phase angle frequency Bode diagram; (<b>d</b>) equivalent circuit diagram.</p>
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<p>The measurement images of Al alloy/steel joints without a pulsed magnetic field during corrosion by a scanning vibrating electrode technique (SVET): (<b>a<sub>1</sub></b>,<b>a<sub>2</sub></b>) soak time—0 h; (<b>b<sub>1</sub></b>,<b>b<sub>2</sub></b>) soak time—1 h; (<b>c<sub>1</sub></b>,<b>c<sub>2</sub></b>) soak time—2 h; (<b>d<sub>1</sub></b>,<b>d<sub>2</sub></b>) soak time—4 h.</p>
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<p>The measurement images of Al alloy/steel joints with a magnetic field strength of 60 mT during corrosion by scanning vibrating electrode technique (SVET): (<b>a<sub>1</sub></b>,<b>a<sub>2</sub></b>) soak time—0 h; (<b>b<sub>1</sub></b>,<b>b<sub>2</sub></b>) soak time—1 h; (<b>c<sub>1</sub></b>,<b>c<sub>2</sub></b>) soak time—2 h; (<b>d<sub>1</sub></b>,<b>d<sub>2</sub></b>) soak time—4 h.</p>
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<p>XPS total spectrum on the passivation film of joints: (<b>a</b>) without magnetic field; (<b>b</b>) with pulsed magnetic field of 60 mT.</p>
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<p>Fitting results of Fe2p orbit of different joints: (<b>a</b>) without magnetic field; (<b>b</b>) with pulsed magnetic field of 60 mT.</p>
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<p>Fitting results of Al2p orbit of different joints: (<b>a</b>) without magnetic field; (<b>b</b>) with pulsed magnetic field of 60 mT.</p>
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<p>Fitting results of Si2p orbit of different joints: (<b>a</b>) without magnetic field; (<b>b</b>) with pulsed magnetic field of 60 mT.</p>
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13 pages, 1729 KiB  
Article
Reducing Avalanche Build-Up Time by Integrating a Single-Photon Avalanche Diode with a BiCMOS Gating Circuit
by Bernhard Goll, Mehran Saadi Nejad, Kerstin Schneider-Hornstein and Horst Zimmermann
Sensors 2024, 24(23), 7598; https://doi.org/10.3390/s24237598 - 28 Nov 2024
Viewed by 319
Abstract
It is shown that the integration of a single-photon avalanche diode (SPAD) together with a BiCMOS gating circuit on one chip reduces the parasitic capacitance a lot and therefore reduces the avalanche build-up time. The capacitance of two bondpads, which are necessary for [...] Read more.
It is shown that the integration of a single-photon avalanche diode (SPAD) together with a BiCMOS gating circuit on one chip reduces the parasitic capacitance a lot and therefore reduces the avalanche build-up time. The capacitance of two bondpads, which are necessary for the connection of an SPAD chip and a gating chip, are eliminated by the integration. The gating voltage transients of the SPAD are measured using an integrated mini-pad and a picoprobe. Furthermore, the gating voltage transients of a CMOS gating circuit and of the BiCMOS gating circuit are compared for the same integrated SPAD. The extension of the 0.35 μm CMOS process by an NPN transistor process module enabled the BiCMOS gating circuit. The avalanche build-up time of the SPAD is reduced to 1.6 ns due to the integration compared to about 3 ns for a wire-bonded off-chip SPAD using the same n+ and p-well as well as the same 0.35 μm technology. Full article
(This article belongs to the Special Issue Advanced CMOS Integrated Circuit Design and Application III)
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Figure 1
<p>Circuit diagram of BiCMOS receiver with integrated SPAD.</p>
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<p>Circuit diagram of CMOS gating circuit.</p>
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<p>Post-layout simulated transients of 6.6 V gating pulses.</p>
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<p>Chip photo of BiCMOS SPAD receiver.</p>
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<p>Schematic cross section of integrated SPAD.</p>
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<p>Measured cathode voltage transient at 500 MHz.</p>
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<p>Measured and overlaid self-discharging transients from dark counts.</p>
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<p>Transients of measured voltage at the CAT node (<b>blue</b>) and derived avalanche current (<b>red</b>) by dependence on time.</p>
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<p>Peak current during avalanche of SPAD vs. substrate voltage.</p>
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<p>Time of occurrence of the avalanche current maximum of the SPAD vs. substrate voltage.</p>
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<p>Extracted mean fall times and their standard deviations over substrate voltage.</p>
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17 pages, 5611 KiB  
Article
Mechanism and Control Strategies for Current Sharing in Multi-Chip Parallel Automotive Power Modules
by Yuqi Jiang, Xuehan Li and Kun Ma
Electronics 2024, 13(23), 4654; https://doi.org/10.3390/electronics13234654 - 25 Nov 2024
Viewed by 400
Abstract
Multi-chip parallel power modules are highly favored in applications requiring high capacity and high switching frequency. However, the dynamic current imbalance between parallel chips caused by asymmetric layouts limits the available capacity. This paper presents a method to optimize dynamic current distribution by [...] Read more.
Multi-chip parallel power modules are highly favored in applications requiring high capacity and high switching frequency. However, the dynamic current imbalance between parallel chips caused by asymmetric layouts limits the available capacity. This paper presents a method to optimize dynamic current distribution by adjusting the lengths and connection points of bond wires. For the first time, a response surface model and nonlinear constraint optimization algorithm are introduced, along with parameter analysis based on finite element methods, to establish the response surface models for the parasitic inductance of bond wires and DBC (direct bonded copper). By leveraging the optimization goals for parasitic inductance and the analytical expressions of all response surfaces, the dynamic current sharing issue was transformed into a nonlinear constrained optimization problem. The solution to this optimization problem identified the optimal connection points for the bond wires, enhancing dynamic current sharing performance. Simulations and experiments were conducted, revealing that the optimized automotive-grade module exhibited a significant reduction in current differences between parallel branches, from 41.7% to 5.03% compared with the original design. This indicated that the proposed optimization scheme for adjusting bond wire connection points could significantly mitigate current disparities, thereby markedly improving current distribution uniformity. Full article
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<p>Topology and physical structure of automotive multi-chip parallel module: (<b>a</b>) topological diagram and (<b>b</b>) physical structure diagram.</p>
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<p>Schematic diagram of test circuit and physical diagram of test platform: (<b>a</b>) schematic diagram of test current and (<b>b</b>) physical diagram of test platform.</p>
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<p>Dynamic current distribution results of two parallel chips in the module at a load current of 600 A: (<b>a</b>) double pulse test waveform, (<b>b</b>) dynamic current distribution of turn-off current for two parallel chips, and (<b>c</b>) dynamic current distribution of turn-on current for two parallel chips.</p>
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<p>Equivalent circuit model for double pulse testing of multi-chip parallel automotive power modules.</p>
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<p>Multi-chip coupled module equivalent circuit chart.</p>
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<p>Single-phase direct bonded copper (DBC) layout of multi-chip coupled power module for vehicle standard.</p>
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<p>Schematic diagram of the bond wire dimension parameters.</p>
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<p>Response surface fitting results and residuals for bond wire inductance and length: (<b>a</b>) response surface fitting results and (<b>b</b>) response surface fitting residuals.</p>
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<p>The layout of the DBC and the location of the chip.</p>
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<p>Parasitic inductance <span class="html-italic">L<sub>cuW</sub></span><sub>1</sub> with <span class="html-italic">x</span><sub>1</sub> and <span class="html-italic">y</span><sub>1</sub> response surface results.</p>
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<p>Parasitic inductance <span class="html-italic">L<sub>cuW</sub></span><sub>2</sub> with <span class="html-italic">x</span><sub>2</sub> and <span class="html-italic">y</span><sub>2</sub> response surface results.</p>
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<p>DBC layout of initial modules and optimized modules: (<b>a</b>) initial modules and (<b>b</b>) optimized modules.</p>
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<p>Measured switching waveforms of two branch circuits in initial and optimized automotive modules: (<b>a</b>) measured turn-on current waveform of the initial module, (<b>b</b>) measured turn-off current waveform of the initial module, (<b>c</b>) measured turn-on current waveform of the optimized module, and (<b>d</b>) measured turn-off current waveform of the optimized module.</p>
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23 pages, 8263 KiB  
Article
Bond–Slip Behavior of High-Strength Stainless Steel Wire Mesh in Engineered Cementitious Composites: Numerical and Theoretical Analysis
by Xuyan Zou, Tao Zhang, Ziyuan Li, Juntao Zhu, Ke Li and Minghao Peng
Materials 2024, 17(23), 5700; https://doi.org/10.3390/ma17235700 - 21 Nov 2024
Viewed by 576
Abstract
This study introduces high-strength non-prestressed steel strands as reinforcement materials into Engineered Cementitious Composites (ECCs) and developed a novel high-strength stainless-steel-strand-mesh (HSSWM)-reinforced ECC with enhanced toughness and corrosion resistance. The bonding performance between HSSWM and an ECC is essential for facilitating effective cooperative [...] Read more.
This study introduces high-strength non-prestressed steel strands as reinforcement materials into Engineered Cementitious Composites (ECCs) and developed a novel high-strength stainless-steel-strand-mesh (HSSWM)-reinforced ECC with enhanced toughness and corrosion resistance. The bonding performance between HSSWM and an ECC is essential for facilitating effective cooperative behavior. The bond behavior between the HSSWM and ECC was investigated through theoretical analysis. A local bond–slip model was proposed based on the average bond–slip model for HSSWM and ECCs. The results indicated that the local bond–slip model provided a more accurate analysis of the bonding performance between HSSWM and the ECC compared to the average bond–slip model. The effects of the ECC’s tensile strength, steel strand diameter, and transverse strand spacing on local bond–slip mechanical behavior were investigated through FEA. The results showed that the local bond–slip model and FE results aligned well with the experimental data. Additionally, the distribution of bond stress between the HSSWM and the ECC was analyzed using the micro-element method based on the local bond–slip model. A prediction model for the critical anchorage length and bond capacity of HSSWM in the ECC was established, and the accuracy of the model was verified. Full article
(This article belongs to the Section Materials Simulation and Design)
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<p>Test setup of HSSWM and the ECC. (<b>a</b>) Test loading setup; (<b>b</b>) diagram of specimens.</p>
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<p>Pull model for HSSWM and ECC (<span class="html-italic">l<sub>d</sub></span> is the spacing of the transverse steel strands).</p>
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<p>Bond stress distribution at different anchorage lengths: (<b>a</b>) short anchorage length; (<b>b</b>) long anchorage length.</p>
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<p>Load–slip curves at specimen ends: (<b>a</b>) H15-h2.4; (<b>b</b>) H18-h2.4; (<b>c</b>) H15-h4.5; (<b>d</b>) H18-h4.5.</p>
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<p>Bond–slip relationship curve between HSSWM and ECCs.</p>
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<p>Schematic diagram of numerical analysis.</p>
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<p>Comparison of load-slip curves for the fitting group: (<b>a</b>) H15-h2.4-30; (<b>b</b>) H15-h4.5-20; (<b>c</b>) H15-h4.5-30; (<b>d</b>) H15-h4.5-40.</p>
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<p>Comparison of load–slip curves for the validation group: (<b>a</b>) H18-h2.4-30; (<b>b</b>) H20-h2.4-30; (<b>c</b>) H18-h4.5-30; (<b>d</b>) H20-h4.5-30; (<b>e</b>) H20-h4.5-30.</p>
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<p>Comparison of load–slip curves for the validation group: (<b>a</b>) H18-h2.4-30; (<b>b</b>) H20-h2.4-30; (<b>c</b>) H18-h4.5-30; (<b>d</b>) H20-h4.5-30; (<b>e</b>) H20-h4.5-30.</p>
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<p>The HX24L element of the ECC. (<b>a</b>) HX24L element of the ECC; (<b>b</b>) translational displacements along the x, y, and z directions at each node; (<b>c</b>) deformations generated by displacements in different directions at each node.</p>
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<p>The 3D nonlinear FE model: (<b>a</b>) HSSWM model; (<b>b</b>) component model.</p>
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<p>Mesh division of 3D nonlinear FE model.</p>
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<p>Tensile constitutive model of the ECC. (<b>a</b>) Based on the total strain multi-segment linear tensile softening model; (<b>b</b>) tensile stress–strain curve for the ECC.</p>
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<p>Compressive constitutive model of the ECC. (<b>a</b>) Based on the total strain multi-segment linear compression model; (<b>b</b>) compressive stress–strain relationship curve of the ECC.</p>
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<p>Fiber concrete model.</p>
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<p>Tensile stress–strain curves of steel strands with different diameters.</p>
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<p>Load–slip curve: (<b>a</b>) H18-h2.4-30; (<b>b</b>) H20-h2.4-30; (<b>c</b>) H15-h4.5-30; (<b>d</b>) H20-h4.5-30; (<b>e</b>) H22-h4.5-30.</p>
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<p>Load displacement curve under different conditions.</p>
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<p>Stress distribution of longitudinal steel strand of specimens with different ECC tensile strengths: (<b>a</b>) <span class="html-italic">f<sub>tc</sub></span> = 1.97 MPa; (<b>b</b>) <span class="html-italic">f<sub>tc</sub></span> = 2.48 MPa; (<b>c</b>) <span class="html-italic">f<sub>tc</sub></span> = 2.83 MP; (<b>d</b>) <span class="html-italic">f<sub>tc</sub></span> = 3.60 MPa.</p>
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<p>Distribution curve of along the anchorage length direction. (<b>a</b>) <span class="html-italic">l<sub>d</sub></span> = 20 mm; (<b>b</b>) <span class="html-italic">l<sub>d</sub></span> = 30 mm; (<b>c</b>) <span class="html-italic">l<sub>d</sub></span> = 20 mm; (<b>d</b>) <span class="html-italic">l<sub>d</sub></span> = 30 mm.</p>
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<p>Bond stress distribution curve along the anchorage length.</p>
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<p>Effect of the spacing between transverse steel tendons on the critical anchorage length.</p>
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16 pages, 5358 KiB  
Article
High-Entropy Alloy Laser Cladding with Cable-Type Welding Wire: Experimental Study and First-Principles Calculations
by Wenjun Wang, Yifei Zheng, Zhihui Cai, Wenjian Zheng, Cai Zhang, Yu Wang, Zhiyong Zhao, Daochen Feng, Yinghe Ma and Jianguo Yang
Metals 2024, 14(11), 1294; https://doi.org/10.3390/met14111294 - 16 Nov 2024
Viewed by 763
Abstract
The Co-Cr-Fe-Ni high-entropy alloy (HEA) is particularly suitable for preparing coatings due to its excellent comprehensive properties. In this study, we use the laser cladding method to prepare Co-Cr-Fe-Ni HEA coatings with Co-Cr-Fe-Ni cable-type welding wire (CTWW) as the filling material and investigated [...] Read more.
The Co-Cr-Fe-Ni high-entropy alloy (HEA) is particularly suitable for preparing coatings due to its excellent comprehensive properties. In this study, we use the laser cladding method to prepare Co-Cr-Fe-Ni HEA coatings with Co-Cr-Fe-Ni cable-type welding wire (CTWW) as the filling material and investigated the dilution rates of the coatings by experimental studies and first-principles calculations. The dilution rate is reduced to about 50% by changing the wire feeding speed, and a Co-Cr-Fe-Ni HEA coating with near nominal composition was prepared by multi-layer cladding. The HEA coating with near nominal composition is successfully prepared in the fourth layer of cladding. The coating is dense and uniform, with good metallurgical bonding. The mechanical properties of the coating were explored using first-principles calculations. All four coatings exhibit a single face-centered cubic (FCC) phase with good mechanical stability in the ground state. The bulk modulus B, shear modulus G, and Young’s modulus E of the four layers of coatings are gradually decreasing from B = 202 GPa, G = 136 GPa, and E = 334 GPa to B = 239 GPa, G = 154 GPa, and E = 380 GPa. The brittleness of the coating shows a trend of first decreasing and then increasing, and the coating closest to the nominal composition has the highest brittleness. Full article
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<p>The laser cladding of the HEA coatings using the Co-Cr-Fe-Ni CTWWs. (<b>a</b>) The stranding wire machine. (<b>b</b>–<b>d</b>) The Co-Cr-Fe-Ni CTWWs. (<b>e</b>) The laser cladding process.</p>
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<p>Fused layer cross-section diagram.</p>
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<p>The unoptimized supercell models of L1, L2, L3, and L4: (<b>a</b>) FCC and (<b>b</b>) BCC.</p>
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<p>A macroscopic image and the heights of HEA multi-layer cladding coatings.</p>
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<p>The cross-sectional metallographic images of four layers of cladding coating samples corroded with 4% nitric acid alcohol. (<b>a</b>) The fusion zone corroded for 5 s. (<b>b</b>) The coating corroded for 5 s. (<b>c</b>) The fusion zone corroded for more than 20 s. (<b>d</b>) The coating corroded for more than 20 s.</p>
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<p>SEM/EDS micrographs of the HEA-based four-layer cladding coating sample: (<b>a</b>) coating and (<b>b</b>) fusion zone.</p>
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<p>The moduli (<span class="html-italic">B</span>, <span class="html-italic">G</span>, and <span class="html-italic">E</span>) of the coatings.</p>
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<p>The Cauchy pressure (<span class="html-italic">C</span><sub>12</sub>–<span class="html-italic">C</span><sub>14</sub>) values of the coatings.</p>
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<p>The Pugh ratio (<span class="html-italic">B</span>/<span class="html-italic">G</span>) and Poisson’s ratio (<span class="html-italic">v</span>) values of the coatings.</p>
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14 pages, 4369 KiB  
Article
Design of High-Speed Thin-Film Lithium Niobate Modulator Utilizing Flip-Chip Bonding with Bump Contacts
by Yihui Yin, Jiayu Yang, Haiou Li, Wanli Yang, Yue Li and Hanyu Li
Electronics 2024, 13(22), 4463; https://doi.org/10.3390/electronics13224463 - 14 Nov 2024
Viewed by 625
Abstract
Currently, the high-speed performance of thin-film lithium niobate electro-optic modulator chips is evolving rapidly. Nevertheless, due to the inherent technical limitations imposed by the packaging design and material architecture, the intrinsic electro-optic bandwidth of thin-film lithium niobate electro-optic modulator chips often exceeds the [...] Read more.
Currently, the high-speed performance of thin-film lithium niobate electro-optic modulator chips is evolving rapidly. Nevertheless, due to the inherent technical limitations imposed by the packaging design and material architecture, the intrinsic electro-optic bandwidth of thin-film lithium niobate electro-optic modulator chips often exceeds the bandwidth of their packaging interfaces, which can constrain the realization of modulation performance. Bump bonding emerges as a high-bandwidth EO interconnection technology, outperforming wire bonding for faster optical communication. In this paper, we present a high-speed thin-film lithium niobate modulator chip tailored for concave–convex bonding, alongside an analysis and design of the chip’s flip-chip bonding packaging. By exploiting the superior electrical characteristics of concave–convex bonding, we effectively mitigate the radio frequency losses of modulator chip and packaging. The simulated half-wave voltage (Vπ) of 3.5 V and E-O modulation bandwidth greater than 150 GHz is obtained for a 0.5 cm long modulator after flip-chip bonding packaging. Full article
(This article belongs to the Special Issue Cognition and Utilization of Electromagnetic Space Signals)
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<p>Overall structure of the High-Speed Thin-Film Lithium Niobate Modulator Utilizing Flip-Chip Bonding with Bump Contacts.</p>
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<p>(<b>a</b>) Schematic diagram of the three-dimensional model for the modulation region of a high-speed thin-film lithium niobate electro-optic modulator; (<b>b</b>) schematic diagram of the single-arm cross-section for the modulation region.</p>
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<p>(<b>a</b>) y-z section view of the flip-chip bonding with bump contacts structure; (<b>b</b>) x–y top view of the flip-chip bonding with bump contacts structure; (<b>c</b>) schematic diagram of the embedded port and bump bonding positions; (<b>d</b>) schematic diagram of the T-shaped electrode structure in the modulation region.</p>
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<p>(<b>a</b>) Embedded port structure and electrical S-parameters obtained by simulation. (<b>b</b>) Linearly tapered port structure and electrical S-parameters obtained by simulation. (<b>c</b>) Curved port structure and electrical S-parameters obtained by simulation.</p>
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<p>(<b>a</b>) Simulation of the TE optical mode in a ridge waveguide. (<b>b</b>) Simulation of the electric field distribution with an electrode spacing of 5 μm.</p>
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<p>(<b>a</b>) The characteristic impedance (left y-axis) and microwave refractive index (right y-axis) of the modulator chip. (<b>b</b>) Simulated RF S<sub>21</sub> and E-O S<sub>21</sub> of 0.5-cm-long modulator chip.</p>
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<p>(<b>a</b>) Three-dimensional diagram of flip-chip bonding with bump contacts. (<b>b</b>) Three-dimensional diagram of wire bonding. (<b>c</b>) The S-parameters for the two bonding methods mentioned above.</p>
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<p>Overall structure of the wire-bonded modulator structure (straight electrode).</p>
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<p>The S-parameter plots for the modulator chip, overall structure of the bump-bonded modulator depicted in <a href="#electronics-13-04463-f002" class="html-fig">Figure 2</a> and the wire-bonded modulator presented in <a href="#electronics-13-04463-f008" class="html-fig">Figure 8</a>.</p>
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14 pages, 11262 KiB  
Article
Effect of Co Addition on the Microstructure and Mechanical Properties of Sn-11Sb-6Cu Babbitt Alloy
by Zhan Cheng, Meng Wang, Bo Wang, Lei Zhang, Ting Zhu, Ningbo Li, Jifa Zhou and Fei Jia
Materials 2024, 17(22), 5494; https://doi.org/10.3390/ma17225494 - 11 Nov 2024
Viewed by 693
Abstract
A Babbitt alloy SnSb11Cu6 with 0–2.0 wt.% Co was synthesized using the induction melting process. This study examined the effect of cobalt (Co) on the microstructure, tensile properties, compressive properties, Brinell hardness, and wear properties of SnSb11Cu6 using optical microscopy (OM), scanning electron [...] Read more.
A Babbitt alloy SnSb11Cu6 with 0–2.0 wt.% Co was synthesized using the induction melting process. This study examined the effect of cobalt (Co) on the microstructure, tensile properties, compressive properties, Brinell hardness, and wear properties of SnSb11Cu6 using optical microscopy (OM), scanning electron microscopy (SEM), energy-dispersive X-ray spectroscopy (EDS), X-ray diffraction (XRD), a universal tensile testing machine, a Brinell hardness tester, and a wear testing machine. The results indicate that the optimal quantity of Co can enhance the microstructure of the Babbitt alloy and promote microstructure uniformity, with presence of Co3Sn2 in the matrix. With the increase in Co content, the tensile and compressive strength of the Babbitt alloy first increased and then decreased, and the Brinell hardness gradually increased with the increase in Co content. The presence of trace Co has a minimal effect on the dry friction coefficient of the Babbitt alloy. When the Co content exceeds 1.5 wt.%, the friction properties of the Babbitt alloy deteriorate significantly. The optimized Babbitt alloy SnSb11Cu6-1.5Co was subsequently fabricated into wires, followed by conducting cold metal transfer (CMT) surfacing experiments. The Co element can promote the growth of interfacial compounds. The microstructure at the interface of the Babbitt alloy/steel is dense, and there is element diffusion between it. The metallurgical bonding is good, and there are serrated compounds relying on the diffusion layer to extend to the direction of the additive layer with serrated compounds extending and growing from the diffusion layer to the additive layer. Overall, Babbitt alloys such as SnSb11Cu6 exhibit improved comprehensive properties when containing 1.5 wt.% Co. Full article
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<p>Casting mold of tensile strength sample.</p>
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<p>OM microstructure of Babbitt alloy SnSb11Cu6 with different Co contents: (<b>a</b>) 0 wt.% Co; (<b>b</b>) 0.5 wt.% Co; (<b>c</b>) 1.0 wt.% Co; (<b>d</b>) 1.5 wt.% Co; (<b>e</b>) 2.0 wt.% Co.</p>
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<p>BSE microstructure of Babbitt alloy SnSb11Cu6 with different Co contents: (<b>a</b>) 0 wt.% Co; (<b>b</b>) 0.5 wt.% Co; (<b>c</b>) 1.0 wt.% Co; (<b>d</b>) 1.5 wt.% Co; (<b>e</b>) 2.0 wt.% Co.</p>
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<p>XRD pattern of Babbitt alloy SnSb11Cu6 with different Co contents (0 wt.%, 1.0 wt.%, 2.0 wt.%).</p>
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<p>Elemental distribution in microstructure of SnSb11Cu6-1.5Co alloy using BEI and EDS: (<b>a</b>) BEI; (<b>b</b>) Sn map; (<b>c</b>) Sb map (<b>d</b>); Cu map; (<b>e</b>) Co map; and (<b>f</b>) combined elemental composition map.</p>
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<p>Co content effect on tensile strength of Babbitt alloy SnSb11Cu6 with 0–2.0 wt.% Co.</p>
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<p>Co content effect on compressive strength of Babbitt alloy SnSb11Cu6 with 0–2.0 wt.% Co.</p>
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<p>Co content effect on Brinell hardness of Babbitt alloy SnSb11Cu6 with 0–2.0 wt.% Co.</p>
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<p>Co content effect on friction coefficient of Babbitt alloy SnSb11Cu6 with 0–2.0 wt.% Co.</p>
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<p>Microstructural interface of Q235 steel/SnSb11Cu6 and Q235 steel/SnSb11Cu6-1.5Co Babbitt alloys: (<b>a</b>,<b>b</b>) OM; (<b>c</b>,<b>d</b>) SEM.</p>
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<p>Elemental distribution analysis of Q235 steel/SnSb11Cu6 and Q235 steel/SnSb11Cu6-1.5Co Babbitt alloys: (<b>a</b>) SnSb11Cu6; (<b>b</b>) SnSb11Cu6-1.5Co.</p>
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12 pages, 2518 KiB  
Article
In Situ Multiphysical Metrology for Photonic Wire Bonding by Two-Photon Polymerization
by Yu Lei, Wentao Sun, Xiaolong Huang, Yan Wang, Jinling Gao, Xiaopei Li, Rulei Xiao and Biwei Deng
Materials 2024, 17(21), 5297; https://doi.org/10.3390/ma17215297 - 31 Oct 2024
Viewed by 627
Abstract
Femtosecond laser two-photon polymerization (TPP) technology, known for its high precision and its ability to fabricate arbitrary 3D structures, has been widely applied in the production of various micro/nano optical devices, achieving significant advancements, particularly in the field of photonic wire bonding (PWB) [...] Read more.
Femtosecond laser two-photon polymerization (TPP) technology, known for its high precision and its ability to fabricate arbitrary 3D structures, has been widely applied in the production of various micro/nano optical devices, achieving significant advancements, particularly in the field of photonic wire bonding (PWB) for optical interconnects. Currently, research on optimizing both the optical loss and production reliability of polymeric photonic wires is still in its early stages. One of the key challenges is that inadequate metrology methods cannot meet the demand for multiphysical measurements in practical scenarios. This study utilizes novel in situ scanning electron microscopy (SEM) to monitor the working PWBs fabricated by TPP technology at the microscale. Optical and mechanical measurements are made simultaneously to evaluate the production qualities and to study the multiphysical coupling effects of PWBs. The results reveal that photonic wires with larger local curvature radii are more prone to plastic failure, while those with smaller local curvature radii recover elastically. Furthermore, larger cross-sectional dimensions contribute dominantly to the improved mechanical robustness. The optical-loss deterioration of the elastically deformed photonic wire is only temporary, and can be fully recovered when the load is removed. After further optimization based on the results of multiphysical metrology, the PWBs fabricated in this work achieve a minimum insertion loss of 0.6 dB. In this study, the multiphysical analysis of PWBs carried out by in situ SEM metrology offers a novel perspective for optimizing the design and performance of microscale polymeric waveguides, which could potentially promote the mass production reliability of TPP technology in the field of chip-level optical interconnection. Full article
(This article belongs to the Special Issue Advances in Laser Processing of Materials)
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<p>Measurement of optical loss and the mechanical properties of PWBs using in situ SEM.</p>
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<p>SEM images of PWBs at increasing TPP-process laser powers; the inserted figure is a magnified view.</p>
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<p>Shows the optical-loss measurements of the PWB. (<b>a</b>) provides a schematic of the photonic wire bond design parameters, highlighting the significant difference between the mode field diameters of a standard single-mode fiber (FiberHome@SMF-G657A1) and the PWB. The typical mode field diameter of an SMF is 10.4 μm, while that of the PWB is less than 3 μm. Simulations and experiments were conducted to evaluate the optical loss for varying bond cross-sectional diameters (w<sub>1</sub>) (<b>b</b>), taper widths (w<sub>2</sub>) (<b>c</b>), taper lengths (L<sub>taper</sub>) (<b>d</b>), and bond curvature radii (R) (<b>e</b>). (<b>f</b>) presents the measured optical losses of printed PWB structures with varying bond cross-sectional diameters in an FA–FA array with a total length of 280 μm, while maintaining the fixed parameters of w<sub>2</sub> = 14 μm, L<sub>taper</sub> = 75 μm, and a curvature radius of R = 40 μm.</p>
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<p>Compression experiments on PWBs with different curvature radii. (<b>a</b>) SEM images and finite-element simulations of the PWB during compression with a curvature radius of 50 μm (<b>a</b>), 40 μm (<b>b</b>), and 30 μm (<b>c</b>). (<b>d</b>,<b>e</b>) Force–displacement curves of the PWB under experimental (EX) and finite-element (FE) simulation conditions. (<b>f</b>,<b>g</b>) Optical loss of the PWB during the compression process under experimental and simulation conditions. The scale bars: 40 μm.</p>
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<p>Compression experiments on PWBs with different bond diameters. (<b>a</b>) This SEM image shows the deformation process of PWB during the compression process; the PWB with a curvature radius of 30 μm exhibits reversible deformation when compressed by 20 μm. (<b>b</b>) Finite-element simulations of the compression response of the PWB structure in Figure (<b>a</b>); (<b>c</b>,<b>d</b>) The force–displacement curve of PWB of under experimental and simulation conditions; (<b>e</b>) The optical loss corresponding to PWB during the compression process. The scale bars: 40 μm.</p>
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14 pages, 15598 KiB  
Article
Properties of Wedge Wire Bonded Connection Between a Composite Copper Core Aluminum Shell Wire and an 18650 Cylindrical Lithium-Ion Battery Cell
by Krzysztof Bieliszczuk and Tomasz M. Chmielewski
Materials 2024, 17(21), 5237; https://doi.org/10.3390/ma17215237 - 28 Oct 2024
Viewed by 585
Abstract
Wedge wire bonding is a solid-state joining process that uses ultrasonic vibrations in combination with compression of the materials to establish an electrical connection. In the battery industry, this process is used to interconnect cylindrical battery cells due to its ease of implementation, [...] Read more.
Wedge wire bonding is a solid-state joining process that uses ultrasonic vibrations in combination with compression of the materials to establish an electrical connection. In the battery industry, this process is used to interconnect cylindrical battery cells due to its ease of implementation, high flexibility and ease of automation. Wire materials typically used in battery pack manufacturing are pure or alloyed aluminum and copper. While copper wires possess better electrical properties, the force used in the bonding process can lead to cell isolator damage and cell thermal runaway. This is an unacceptable result of the bonding process and has led to the development of new types of composite wires containing a copper core embedded in an aluminum shell. This material has the advantage of high copper electrical and thermal conductivity combined with less aggressive bonding parameters of the aluminum wire. The aim of this study was to establish a process window for the wedge wire bonding of 400 µm composite copper–aluminum Heraeus CucorAl Plus wire on the surface of a BAK 18650 battery cell. This study was conducted using a Hesse Bondjet BJ985 CNC wire bonder fitted with an RBK03 bond head designed for the bonding of copper wires. The methods used in this study included light and scanning electron microscopy of bond and battery cell cross-sections, shear testing on the XYZtec Sigma bond tester system, and energy dispersive spectroscopy. The results were compared with a previous study conducted using a wire of the same diameter and made out of high-purity aluminum. Full article
(This article belongs to the Special Issue Advances in Solid-State Welding Processes)
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<p>Wire bonding process parameters.</p>
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<p>Process curves—deformation (green) and transducer voltage (blue)—for parameter sets 1 (<b>a</b>) and 2 (<b>b</b>).</p>
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<p>CucorAl Plus wire transversal (<b>a</b>) and longitudinal (<b>b</b>) cross-section in 200× light microscopy.</p>
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<p>CucorAl Plus wire transversal (<b>a</b>) and longitudinal (<b>b</b>) cross-sections in 1000× light microscopy.</p>
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<p>Measurements of the wire bond cross-section with 200× magnification.</p>
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<p>Transversal cross-section made with parameter sets 1 (<b>a</b>) and 2 (<b>b</b>) in 200× magnification.</p>
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<p>Transversal cross-section made with parameter sets 1 (<b>a</b>) and 2 (<b>b</b>) in 1000× magnification.</p>
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<p>Longitudinal cross-sections made with parameter sets 1 (<b>a</b>) and 2 (<b>b</b>) in 50× magnification.</p>
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<p>SEM imaging of the transversal cross-section of the bond made with bonding parameter set 2 in magnifications of 400× (<b>a</b>), 1000× (<b>b</b>), 5000× with EDS line (<b>c</b>) and 30,000× with EDS line (<b>d</b>).</p>
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<p>Line EDS result of the wire bond—weight %.</p>
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<p>Line EDS result of the wire Cu/Al interface—weight %.</p>
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<p>SEM imaging of the longitudinal cross-section of the bond made with bonding parameter set 2 in magnifications of 100× (<b>a</b>), 500× (<b>b</b>), 1000× (<b>c</b>) and 5000× with EDS line (<b>d</b>).</p>
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<p>Line EDS results for longitudinal cross-section.</p>
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<p>Shear force curves for samples made with bonding parameter set 1.</p>
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<p>Shear z curves for aluminum wire from previous study [<a href="#B16-materials-17-05237" class="html-bibr">16</a>].</p>
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<p>Shear force curves for samples made with bonding parameter set 2.</p>
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<p>Macroscopic images (50×) of the sheared area for bonds made with parameter set 1 (<b>a</b>,<b>c</b>) and parameter set 2 (<b>b</b>,<b>d</b>).</p>
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23 pages, 2933 KiB  
Article
Shear Bond Strength in Stone-Clad Façades: Effect of Polypropylene Fibers, Curing, and Mechanical Anchorage
by Vahid Shafaie, Oveys Ghodousian, Amin Ghodousian, Mohammad Gorji, Hossein Mehdikhani and Majid Movahedi Rad
Polymers 2024, 16(21), 2975; https://doi.org/10.3390/polym16212975 - 24 Oct 2024
Viewed by 508
Abstract
This study investigates the shear bond strength between four widely used façade stones—travertine, granite, marble, and crystalline marble—and concrete substrates, with a particular focus on the role of polypropylene fibers in adhesive mortars. The research evaluates the effects of curing duration, fiber dosage, [...] Read more.
This study investigates the shear bond strength between four widely used façade stones—travertine, granite, marble, and crystalline marble—and concrete substrates, with a particular focus on the role of polypropylene fibers in adhesive mortars. The research evaluates the effects of curing duration, fiber dosage, and mechanical anchorage on bond strength. Results demonstrate that Z-type anchorage provided the highest bond strength, followed by butterfly-type and wire tie systems. Extended curing had a significant impact on bond strength for specimens without anchorage, particularly for travertine. The incorporation of polypropylene fibers at 0.2% volume in adhesive mortar yielded the strongest bond, although lower and higher dosages also positively impacted the bonding. Furthermore, the study introduces a novel fuzzy logic model using the Dombi family of t-norms, which outperformed linear regression in predicting bond strength, achieving an R2 of up to 0.9584. This research emphasizes the importance of optimizing fiber dosage in adhesive mortars. It proposes an advanced predictive model that could enhance the design and safety of stone-clad façades, offering valuable insights for future applications in construction materials. Full article
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<p>Research flowchart: experimental design and predictive modeling.</p>
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<p>The polypropylene fibers.</p>
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<p>Schematic back anchor used in the study: (<b>a</b>) butterfly-type clip, (<b>b</b>) -type clip.</p>
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<p>Façade stones used in the study: (<b>a</b>) travertine, (<b>b</b>) granite, (<b>c</b>) marble, (<b>d</b>) crystalline marble.</p>
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<p>Anchorage installations on façade stones: (<b>a</b>) butterfly-type clip, (<b>b</b>) Z-type clip, (<b>c</b>) wire tie, (<b>d</b>) schematic of composite specimen showing three layers: 15-cm concrete substrate, 3-cm overlay of adhesive mortar, and 2-cm façade stone, with schematic anchorage placement, (<b>e</b>) schematic installation of wire tie.</p>
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<p>Preparation of composite specimens; (<b>a</b>) laboratory preparation process: left—placement of anchorage, middle—molding of the specimen, right—application of adhesive mortar; (<b>b</b>) schematic representation of the final assembly.</p>
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<p>Compressive strength test setup for 5 cm cubic adhesive mortar specimens.</p>
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<p>Shear splitting test: (<b>a</b>) composite specimen before failure, (<b>b</b>) composite specimen after failure at the adhesive interface, and (<b>c</b>) schematic representation of the shear splitting test setup.</p>
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<p>Shear bond strength between façade stones and concrete substrate for (<b>a</b>) travertine, (<b>b</b>) granite, (<b>c</b>) marble, and (<b>d</b>) crystalline marble after 7-day and 28-day curing periods with varying polypropylene fiber content (Ctrl, P1, P2, P3) and anchorage types (No anchorage, Z-type clip, butterfly-type clip, and wire tie).</p>
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<p>Bond strength growth between façade stones and concrete substrates due to the presence of different types of anchorage systems (Z-type, butterfly-type, and wire tie), compared to the no-anchorage condition. The graph reflects the results based on 7-day curing specimens for various façade stones, including travertine, granite, marble, and crystalline marble.</p>
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<p>Bond strength changes (%) due to the increase in curing duration from 7 days to 28 days for various façade stones and anchorage types.</p>
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<p>Shear bond strength for four façade stones across polypropylene fiber dosages averaged over all four anchorage conditions.</p>
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<p>Linear regression predictions compared with experimental results.</p>
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<p>Membership functions for three inputs: (<b>a</b>) anchorage type, (<b>b</b>) water absorption of the façade stone, and (<b>c</b>) volume percentage of polypropylene fiber.</p>
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<p>Prediction of bond strength with proposed fuzzy logic inference system using Dombi family of t-norms: (<b>a</b>) λ = 0.001, (<b>b</b>) λ = 0.01, (<b>c</b>) λ = 0.05, (<b>d</b>) λ = 0.1, (<b>e</b>) λ = 0.25, (<b>f</b>) λ = 0.5, and (<b>g</b>) λ = 1.</p>
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<p>Prediction of bond strength with proposed fuzzy logic inference system using Dombi family of t-norms: (<b>a</b>) λ = 5, (<b>b</b>) λ = 10, (<b>c</b>) λ = 50, and (<b>d</b>) λ = 100.</p>
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<p>Prediction of bond strength with proposed fuzzy logic inference system using Dombi family of t-norms: (<b>a</b>) λ = 200, (<b>b</b>) λ = 300, (<b>c</b>) λ = 400, (<b>d</b>) λ = 500, (<b>e</b>) λ = 600, (<b>f</b>) λ = 700, (<b>g</b>) λ = 800, (<b>h</b>) λ = 900, and (<b>i</b>) λ = 1000.</p>
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23 pages, 8234 KiB  
Article
Bond Strength and Corrosion Protection Properties of Hot-Dip Galvanized Prestressing Reinforcement in Normal-Strength Concrete
by Petr Pokorný, Tomáš Chobotský, Nikola Prodanovic, Veronika Steinerová and Karel Hurtig
J. Compos. Sci. 2024, 8(10), 407; https://doi.org/10.3390/jcs8100407 - 4 Oct 2024
Cited by 2 | Viewed by 793
Abstract
Several prestressing reinforced structures have recently collapsed due to chloride-induced steel corrosion. This study investigates the effect of the corrosion of hot-dip galvanized conventional prestressing steel reinforcement under hydrogen evolution on bond strength in normal-strength concrete. The impact of hydrogen evolution on the [...] Read more.
Several prestressing reinforced structures have recently collapsed due to chloride-induced steel corrosion. This study investigates the effect of the corrosion of hot-dip galvanized conventional prestressing steel reinforcement under hydrogen evolution on bond strength in normal-strength concrete. The impact of hydrogen evolution on the porosity of cement paste at the interfacial transition zone (ITZ) is verified through image analysis. The whole surface of prestressing strands is hot-dip galvanized, and their corrosion behavior when embedded in the cement paste is investigated by measuring the time dependence of the open-circuit potential. Concerning the uniformity of the hot-dip galvanized coating and its composition, it is advisable to coat the individual wires of the prestressing reinforcement and subsequently form a strand. It is demonstrated that the corrosion of the coating under the evolution of hydrogen in the cement paste reduces the bond strength of hot-dip galvanized reinforcement in normal-strength concrete. Image analysis after 28 days of cement paste aging indicates insignificant filling of hydrogen-generated pores by zinc corrosion products. Applying an additional surface treatment (topcoat) stable in an alkaline environment is necessary to avoid corrosion of the coating under hydrogen evolution and limit the risk of bond strength reduction. Full article
(This article belongs to the Section Composites Manufacturing and Processing)
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<p>Example of corrosion damage of conventional prestressing steel reinforcement stimulated by chloride anions—archive of Klokner Institute of CTU in Prague.</p>
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<p>Collapse of Troja footbridge in Prague caused by corrosion of conventional prestressing steel reinforcement—archive of Klokner Institute of CTU in Prague.</p>
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<p>Typical composition of hot-dip galvanized coating on low-silicon steel (outside the so-called Sandelin area and the area with significant silicon content in steel)—reprinted from [<a href="#B25-jcs-08-00407" class="html-bibr">25</a>].</p>
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<p>Modelling of sample for measurement E<sub>corr</sub>/FeCr18Ni9 of hot-dip galvanized wire (from prestressing strand) in cement paste: (<b>A</b>) Modelling of sample for measurement; (<b>B</b>) Real view of sample for measurement.</p>
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<p>Figure (<b>A</b>) shows how the steel strand was positioned in the center of the concrete cube using a wooden wedge and a clamp. Figure (<b>B</b>) shows a group of concrete cubes prepared for the pull-out test.</p>
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<p>The pull-out bond strength test between a strand sample and normal-strength concrete: (<b>A</b>) experimental setup schema; (<b>B</b>) the setup of MTS 500 kN loading machine.</p>
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<p>Cement paste specimens were prepared using both types of prestressing steel to determine the porosity of the cement paste at the interface.</p>
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<p>Evaluation of coating formation on individual prestressing wires by optical microscopy – overview.</p>
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<p>Evaluation of continuity of HDG coating between (<b>A</b>) the outer wires of prestressing steel strand—cracks in the coating; (<b>B</b>) the outer and inner wire of prestressing steel strand—presence of discontinuous coating.</p>
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<p>HDG coating on the surface of the prestressing strand—detailed view: (<b>A</b>) The surface of the outer wire; (<b>B</b>) The surface of the inner wire.</p>
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<p>Comparison of surface roughness (R<sub>a</sub>) between uncoated steel (US) and hot-dip galvanized (HDG) prestressing steel wires.</p>
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<p>Time development (250 h of testing) of open-circuit potential for samples: (<b>A</b>) the uncoated (US) prestressing steel in cement paste; (<b>B</b>) HDG prestressing steel in cement paste.</p>
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<p>Time development (250 h of testing) of open-circuit potential for samples: (<b>A</b>) the uncoated (US) prestressing steel in cement paste; (<b>B</b>) HDG prestressing steel in cement paste.</p>
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<p>Results of bond strength tests: (<b>A</b>) bond stress–slip curves overview (curves—mean values, error bars—standard deviation); (<b>B</b>) bond stress–slip curves for low slip values (curves—mean values, error bars—standard deviation) after 28 days of concrete aging.</p>
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<p>Failure pattern after pull-out test for concrete samples with: (<b>A</b>) uncoated (US) prestressing steel; (<b>B</b>) HDG prestressing steel (after 28 days of concrete aging).</p>
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<p>Hot-dip galvanized coating cracks/fracture pattern on surface of prestressing steel after pull-out test: (<b>A</b>) fracture pattern in η phase; (<b>B</b>) fracture pattern in ζ phase (FeZn<sub>13</sub>) after 28 days of concrete aging.</p>
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<p>Interfacial transition zone of uncoated prestressing steel in cement paste after 28 days of aging (overview) and SEM image of detailed porous structure of cement paste from ITZ.</p>
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<p>Interfacial transition zone of hot-dip galvanized prestressing steel in cement paste after 28 days of aging (overview) and SEM image of detailed porous structure of cement paste from ITZ.</p>
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<p>Bar chart of the total area of pores of cement paste taken from different interfacial transition zones (ITZs).</p>
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<p>Diffraction pattern of corrosion products precipitated on surface of hot-dip galvanized prestressing steel after 28 days of curing in cement paste.</p>
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17 pages, 6901 KiB  
Article
Analysis of Power Modules Including Phase Change Materials in the Top Interconnection of Semiconductor Devices
by Rabih Khazaka, Rachelle Hanna, Yvan Avenas and Stephane Azzopardi
Electron. Mater. 2024, 5(4), 204-220; https://doi.org/10.3390/electronicmat5040014 - 1 Oct 2024
Viewed by 823
Abstract
Power modules can occasionally be exposed to brief power peaks, causing overheating and premature failure of the power semiconductor devices. In order to overcome this issue without oversizing the module or its cooling system, this study aims to design a new class of [...] Read more.
Power modules can occasionally be exposed to brief power peaks, causing overheating and premature failure of the power semiconductor devices. In order to overcome this issue without oversizing the module or its cooling system, this study aims to design a new class of power modules with integrated Phase Change Material (PCM) in a container serving as a top device interconnection. Simulations and experiments are performed with two organic PCMs, and the interest in adding copper foam is discussed. Under various test conditions, the results show that the simulations agree well with the experiments. Hence, virtual prototyping can be very useful for sizing containers based on a specific mission profile. For a constant selected PCM volume (around 1 cm3/device) and with a convection heat transfer coefficient value of 800 W.m−2.K−1, the solution allows achieving a junction temperature reduction of about 35 °C (erythritol and 90% porosity copper foam) compared to a wire-bonded conventional technique. Repetitive power cycles can be achieved with both materials, but the selection of the PCM should be conducted cautiously based on the mission profile. The two selected organic PCMs show degradation of their latent heat of fusion and mass loss during high-temperature isothermal aging in air above 130 °C. By assuming as endpoint criterion the reduction of energy storage by 50% compared to the initial state, the lifetime of erythritol and RT100 is evaluated to be about 100 and 340 h, respectively, during aging at 150 °C. Full article
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<p>(<b>a</b>) 3D schematic view of a phase-leg power module using the proposed solution. Two SiC MOSFETs are used for each switch. (<b>b</b>) Cross-section view showing 1 SiC MOSFET.</p>
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<p>(<b>a</b>) Top view of the assembly and (<b>b</b>) the device under thermal testing.</p>
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<p>3D view and cross-section view of the simulated geometry, including two MOSFETs in parallel.</p>
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<p>DSC measurements during heating and cooling of RT100 (<b>a</b>) and erythritol (<b>b</b>) at a rate of 10 °C/min. The measured density and thermal conductivity used in the simulation are shown in the graphs.</p>
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<p>(<b>a</b>) Simulated (lines) and measured (markers) <span class="html-italic">Tj</span> as a function of time for an empty container, a container filled with RT100, and a container filled with RT100 and Cu foam. (<b>b</b>) Calculated PCM melt volume ratio as a function of time for a container filled with RT100 and a container filled with RT100 and Cu foam.</p>
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<p>(<b>a</b>) Images of the top of the container and simulated temperature of the PCM alone after 1 s, 30 s, 60 s and 120 s of power peak application. (<b>b</b>) Images of the top of the container and simulated temperature of the PCM with copper foam after 1 s, 5 s, 10 s and 30 s of power peak application.</p>
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<p>Simulated (lines) and measured (markers) <span class="html-italic">Tj</span> as a function of time under two peak power values for containers with copper foam before and after filling with erythritol. The black line represents the simulated data for the wire-bonded conventional structure added for comparison.</p>
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<p>(<b>a</b>) Power profiles applied during preliminary tests and (<b>b</b>) power profiles used for repetitive power cycling.</p>
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<p>(<b>a</b>) Measured <span class="html-italic">Tj</span> as a function of time for different power cycles for RT100 with copper foam, and (<b>b</b>) measured <span class="html-italic">Tj</span> as a function of time during the heating phase for erythritol with copper foam.</p>
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<p>DSC measurements before and after 100 h of isothermal aging at 130 °C, 140 °C, and 150 °C for RT100 (<b>a</b>) and erythritol (<b>b</b>).</p>
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<p>(<b>a</b>) Color change of tested PCMs after 100h of aging and (<b>b</b>) mass loss ratio at various aging temperatures.</p>
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<p>ln(<span class="html-italic">L<sub>t</sub></span>/<span class="html-italic">L</span><sub>0</sub>) vs. aging time at various aging temperatures for RT100 (<b>a</b>) and erythritol (<b>b</b>).</p>
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<p>Arrhenius plot of the reaction rate constant for erythritol and RT100.</p>
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<p>Calculated latent heat of fusion as a function of aging time at different temperatures between 130 °C and 160 °C for RT100 (<b>a</b>) and erythritol (<b>b</b>).·</p>
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<p>Calculated storage energy ratio as a function of aging time at 150 °C in air.</p>
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21 pages, 6102 KiB  
Article
Optimization of MOSFET Copper Clip to Enhance Thermal Management Using Kriging Surrogate Model and Genetic Algorithm
by Yubin Cheon, Jaehyun Jung, Daeyeon Ki, Salman Khalid and Heung Soo Kim
Mathematics 2024, 12(18), 2949; https://doi.org/10.3390/math12182949 - 22 Sep 2024
Viewed by 1023
Abstract
Metal–oxide–semiconductor field-effect transistors (MOSFETs) are critical in power electronic modules due to their high-power density and rapid switching capabilities. Therefore, effective thermal management is crucial for ensuring reliability and superior performance. This study used finite element analysis (FEA) to evaluate the electro-thermal behavior [...] Read more.
Metal–oxide–semiconductor field-effect transistors (MOSFETs) are critical in power electronic modules due to their high-power density and rapid switching capabilities. Therefore, effective thermal management is crucial for ensuring reliability and superior performance. This study used finite element analysis (FEA) to evaluate the electro-thermal behavior of MOSFETs with copper clip bonding, showing a significant improvement over aluminum wire bonding. The aluminum wire model reached a maximum temperature of 102.8 °C, while the copper clip reduced this to 74.6 °C. To further optimize the thermal performance, Latin Hypercube Sampling (LHS) generated diverse design points. The FEA results were used to select the Kriging regression model, chosen for its superior accuracy (MSE = 0.036, R2 = 0.997, adjusted R2 = 0.997). The Kriging model was integrated with a Genetic Algorithm (GA), further reducing the maximum temperature to 71.5 °C, a 4.20% improvement over the original copper clip design and a 43.8% reduction compared to aluminum wire bonding. This integration of Kriging and the GA to the MOSFET copper clip package led to a significant improvement in the heat dissipation and overall thermal performance of the MOSFET package, while also reducing the computational power requirements, providing a reliable and efficient solution for the optimization of MOSFET copper clip packages. Full article
(This article belongs to the Section Engineering Mathematics)
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<p>Picture of the TO247 MOSFET package: (<b>a</b>) aluminum wire bonding package; (<b>b</b>) copper clip bonding package.</p>
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<p>Comprehensive optimization methodology for copper clip size in MOSFETs using FEA, LHS, and a Kriging surrogate model, assisted by a GA.</p>
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<p>TO247 MOSFET package models: (<b>a</b>) TO247 MOSFET package with encapsulation; (<b>b</b>) aluminum wire bonded TO247 MOSFET package; (<b>c</b>) copper clip bonded TO247 MOSFET package. The colors indicate key components of the MOSFET package: light green represents the die, deep green represents the drain terminal, blue denotes the source terminal, and red indicates the gate terminal.</p>
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<p>Applied boundary conditions for the electro−thermal analysis of MOSFET packages; (A) Lead frame, (B) Gate terminal, (C) Source terminal, (D) Drain terminal, (E) Encapsulation.</p>
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<p>Mesh generated for TO247 MOSFET package models: (<b>a</b>) mesh generated for TO247 MOSFET package with encapsulation; (<b>b</b>) mesh generated for copper clip bonded TO247 MOSFET package; (<b>c</b>) mesh generated for aluminum wire bonded TO247 MOSFET package.</p>
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<p>Parameters defined for the DoEs: (<b>a</b>) horizontal length from the edge to the clip; (<b>b</b>) width of the clip; (<b>c</b>) thickness of the clip.</p>
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<p>Mesh convergence plots for MOSFET thermal analysis: (<b>a</b>) mesh convergence for aluminum wire MOSFET; (<b>b</b>) mesh convergence for copper clip MOSFET.</p>
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<p>Thermal analysis results showing temperature distribution: (<b>a</b>) Al wire bonding and (<b>b</b>) copper clip bonding configurations.</p>
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<p>A 3D scatter plot of the LHS results for the copper clip parameters, where the <span class="html-italic">x</span>-axis represents the parameters. The <span class="html-italic">x</span>-axis, <span class="html-italic">y</span>-axis and <span class="html-italic">z</span>-axis represents the values of parameter a (horizontal length from the die edge to the clip), parameter c (clip thickness), parameter b (clip width), respectively.</p>
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<p>Comparison of the predicted versus the simulated temperatures for various surrogate models: (<b>a</b>) linear regression; (<b>b</b>) polynomial regression; (<b>c</b>) Kriging regression; (<b>d</b>) RBF regression; (<b>e</b>) support vector regression; (<b>f</b>) neural network.</p>
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19 pages, 15534 KiB  
Article
Research on Gate Charge Degradation of Multi-Chip IGBT Modules in Power Supply for Unmanned Aerial Vehicles
by Yuheng Li, Zhiquan Zhou, Jinlong Wang, Lina Wang and Chenxu Wang
Electronics 2024, 13(18), 3664; https://doi.org/10.3390/electronics13183664 - 14 Sep 2024
Viewed by 577
Abstract
In recent years, with the burgeoning application of high voltage in various industrial sectors, the deployment of unmanned equipment, such as industrial heavy-load Unmanned Aerial Vehicles (UAVs), incorporating high-capacity Insulated-Gate Bipolar Transistors (IGBTs), has become increasingly prevalent. The demand for high-voltage IGBT modules [...] Read more.
In recent years, with the burgeoning application of high voltage in various industrial sectors, the deployment of unmanned equipment, such as industrial heavy-load Unmanned Aerial Vehicles (UAVs), incorporating high-capacity Insulated-Gate Bipolar Transistors (IGBTs), has become increasingly prevalent. The demand for high-voltage IGBT modules in UAV is continuously growing; therefore, exploring methods to predict fault precursor parameters of multi-chip IGBT modules is crucial for the operational health management of unmanned equipment like UAVs. This paper analyzes the gate charge degradation in multi-chip IGBT modules after thermal cycling, which can be used to evaluate the operational state of these modules. Furthermore, to delve into the electrical response of a gate drive circuit caused by local damage within the IGBT module, an RLC model incorporating parasitic parameters of the gate drive circuit is established, and a sensitivity analysis of the peak current in the gate charge circuit is provided. Additionally, in the experimental circuit, an open sample of an IGBT module with partial bond wires lifted off is used to simulate actual faults. The analysis and experimental results indicate that the peak current of the gate charge is closely related to L and C. The significant deviation in the gate current, influenced by the partial bond wires lift-off, can provide a basis for the development of predictive methods for IGBT modules. Full article
(This article belongs to the Section Industrial Electronics)
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<p>UAV equipped with IGBT modules.</p>
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<p>UAV motor signal path.</p>
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<p>Cross-section of IGBT unit cell.</p>
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<p>Cross-section of IGBT module. Reprinted with permission from ref. [<a href="#B30-electronics-13-03664" class="html-bibr">30</a>]. 1999–2024 John Wiley &amp; Sons, Inc. or related companies.</p>
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<p>Cross-section of IGBT module.</p>
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<p>Stray inductances of bond wires in IGBT module.</p>
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<p>Equivalent circuit of IGBT module. Reprinted with permission from ref. [<a href="#B37-electronics-13-03664" class="html-bibr">37</a>]. 2019 Springer Nature Switzerland AG.</p>
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<p>Typical gate charge.</p>
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<p>Waveforms of the gate voltage at different <math display="inline"><semantics> <msub> <mi>V</mi> <mrow> <mi>C</mi> <mi>E</mi> </mrow> </msub> </semantics></math>.</p>
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<p>Waveforms of the gate current at different <math display="inline"><semantics> <msub> <mi>V</mi> <mrow> <mi>C</mi> <mi>E</mi> </mrow> </msub> </semantics></math>.</p>
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<p>Sensitivity analysis of L and C.</p>
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<p>Experimental setup circuit (a).</p>
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<p>Experimental setup circuit (b).</p>
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<p>Gate charge current pre- and post- partial bond wires lift-off.</p>
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<p>Gate charge current versus <math display="inline"><semantics> <msub> <mi>V</mi> <mrow> <mi>C</mi> <mi>E</mi> </mrow> </msub> </semantics></math> post-six bond wires lift-off.</p>
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<p>Changes of gate voltage and current when <math display="inline"><semantics> <msub> <mi>V</mi> <mrow> <mi>C</mi> <mi>E</mi> </mrow> </msub> </semantics></math> = 200 V: (<b>a</b>) Description of waveform before the defect occurs. (<b>b</b>) Description of the waveform after the defect occurs. The yellow line is <math display="inline"><semantics> <msub> <mi>V</mi> <mrow> <mi>C</mi> <mi>E</mi> </mrow> </msub> </semantics></math>, the blue line is <math display="inline"><semantics> <msub> <mi>V</mi> <mrow> <mi>G</mi> <mi>E</mi> </mrow> </msub> </semantics></math>, and the red line is <math display="inline"><semantics> <msub> <mi>i</mi> <mi>g</mi> </msub> </semantics></math>.</p>
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<p>Changes of gate voltage and current when <math display="inline"><semantics> <msub> <mi>V</mi> <mrow> <mi>C</mi> <mi>E</mi> </mrow> </msub> </semantics></math> = 420 V: (<b>a</b>) Description of waveform before the defect occurs. (<b>b</b>) Description of the waveform after the defect occurs. The yellow line is <math display="inline"><semantics> <msub> <mi>V</mi> <mrow> <mi>C</mi> <mi>E</mi> </mrow> </msub> </semantics></math>, the blue line is <math display="inline"><semantics> <msub> <mi>V</mi> <mrow> <mi>G</mi> <mi>E</mi> </mrow> </msub> </semantics></math>, and the red line is <math display="inline"><semantics> <msub> <mi>i</mi> <mi>g</mi> </msub> </semantics></math>.</p>
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<p>Turn-on waveforms of the gate voltage according to different temperatures.</p>
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