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Topic Editors

Department of Metallurgy and Recykling, Silesian University of Technology, Krasińskiego 8, 40-019 Katowice, Poland
Environmental Research Department, Faculty of Technology The Institute of Technology and Business in České Budějovice, Okružní 517/10, 370 01 České Budějovice, Czech Republic

Advanced Processes in Metallurgical Technologies

Abstract submission deadline
closed (31 October 2023)
Manuscript submission deadline
closed (31 December 2023)
Viewed by
85667

Topic Information

Dear Colleagues,

The aim of this collection of articles is to contribute to the development of metallurgy and materials engineering from the point of view of a multidisciplinary approach, taking into account the issues of environmental protection, generated waste, disposal options, and energy savings. Currently, the production of metals is a very important problem, especially with their high prices and enormous demand. This problem needs to be looked at from both the economic side, i.e., developing technologies that need less energy, and the ecological one, i.e., developing technologies that emit fewer pollutants and at the same time promote green technologies. The aim of this article collection is to contribute to advancements in the fields of metallurgy and engineering materials. We invite researchers from these and other fields to publish interesting articles on this topic to integrate interdisciplinary research gathering experimental and theoretical input. Topics of interest could include but are not limited to:

  • Science and technology of materials; 
  • Metals recovery from waste; 
  • Energy-saving processes; 
  • Physics and numerical modeling; 
  • Advanced metallurgical technologies;
  • Reduction of pollutant emissions;
  • Numerical simulation.

Prof. Dr. Mariola Saternus
Dr. Ladislav Socha
Topic Editors

Keywords

  • metallurgy
  • steel industry
  • waste treatment
  • energy problems
  • coatings
  • aluminum refining
  • zinc recovery
  • functional materials
  • modeling
  • metals recovery
  • artificial intelligence
  • corrosion resistance
  • welding processes

Participating Journals

Journal Name Impact Factor CiteScore Launched Year First Decision (median) APC
Applied Sciences
applsci
2.5 5.3 2011 18.4 Days CHF 2400
Energies
energies
3.0 6.2 2008 16.8 Days CHF 2600
Journal of Manufacturing and Materials Processing
jmmp
3.3 5.1 2017 16.5 Days CHF 1800
Materials
materials
3.1 5.8 2008 13.9 Days CHF 2600
Metals
metals
2.6 4.9 2011 17.8 Days CHF 2600

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Published Papers (39 papers)

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13 pages, 227 KiB  
Editorial
Advanced Processes in Metallurgical Technologies
by Mariola Saternus and Ladislav Socha
Metals 2024, 14(9), 1049; https://doi.org/10.3390/met14091049 - 14 Sep 2024
Viewed by 976
Abstract
The production of metals and their alloys will continue to increase in the coming years, mainly due to the growing demand for these products [...] Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
14 pages, 3337 KiB  
Article
Study of Assimilation of Cored Wire into Liquid Steel Baths
by Edgar-Ivan Castro-Cedeno, Julien Jourdan, Jonathan Martens, Jean-Pierre Bellot and Alain Jardy
Metals 2024, 14(4), 462; https://doi.org/10.3390/met14040462 - 15 Apr 2024
Cited by 1 | Viewed by 1826
Abstract
Cored wire is a widespread technology used for performing additions into liquid metal baths as an alternative to bulk-additions. A laboratory-scale study was performed in which the kinetics of assimilation of cored wire in liquid steel baths were studied. An original dataset of [...] Read more.
Cored wire is a widespread technology used for performing additions into liquid metal baths as an alternative to bulk-additions. A laboratory-scale study was performed in which the kinetics of assimilation of cored wire in liquid steel baths were studied. An original dataset of positions of the wire/melt interface of cored wire as a function of the time and steel bath temperature was produced. The dataset was compared against results of simulations made with a transient 1D (radial) thermal model of the assimilation of cored wire, and demonstrated reasonable agreement. Hence, this paper provides a dataset that can be used as a resource for the validation of future developments in the field of modeling cored wire injection into liquid metal baths. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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Graphical abstract

Graphical abstract
Full article ">Figure 1
<p>Schematic representation of possible assimilation routes for cored wires immersed into liquid steel.</p>
Full article ">Figure 2
<p>Schematic of experimental apparatus used for the experiments of the immersion of cored wire into liquid steel baths.</p>
Full article ">Figure 3
<p>Photograph of extraction of wire V from the bath during the immersion experiment. Notice the shell of solidified bath material around the cored wire.</p>
Full article ">Figure 4
<p>Cored wires with solidified metal shell after wire immersion experiments at various initial steel bath temperatures: (<b>a</b>) wire I (1818 K); (<b>b</b>) wire II (1823 K); (<b>c</b>) wire III (1823 K); (<b>d</b>) wire IV (1853 K); (<b>e</b>) wire V (1858 K); (<b>f</b>) wire VI (1873 K).</p>
Full article ">Figure 5
<p>Schematic of the wire + solidified shell ensemble post-immersion. The immersion times of radial sections throughout the length of the wire can be estimated using Equations (1)–(4), using the measured length of wire before immersion (<span class="html-italic">L<sub>bef</sub></span>), length of wire after immersion (<span class="html-italic">L<sub>aft</sub></span>), length of non-immersed part of the wire (<span class="html-italic">L<sub>nim</sub></span>), and timestamps of camera recordings (<span class="html-italic">t<sub>im</sub></span>, <span class="html-italic">t<sub>hold</sub></span>, <span class="html-italic">t<sub>ext</sub></span>).</p>
Full article ">Figure 6
<p>Schematic of the procedure for measuring the diameter of the wire + shell ensemble post-immersion. To obtain estimations of mean diameter, the specimen was rotated for each of the 10 individual measurements taken at fixed wire lengths.</p>
Full article ">Figure 7
<p>Comparison of estimations of radius of the wire + shell ensemble (experimental) and computed evolution of the wire/melt interface (simulations): (<b>a</b>) wire <span class="html-italic">I</span> (1818 K); (<b>b</b>) wire <span class="html-italic">II</span> (1823 K), (<b>c</b>) wire <span class="html-italic">III</span> (1823 K); (<b>d</b>) wire <span class="html-italic">IV</span> (1853 K); (<b>e</b>) wire <span class="html-italic">V</span> (1858 K); (<b>f</b>) wire <span class="html-italic">VI</span> (1873 K).</p>
Full article ">Figure 8
<p>Deviations between estimations of thickness of the casing + shell ensemble (experimental) and thicknesses from the computed evolution of the wire/melt interface (simulations).</p>
Full article ">Figure 9
<p>Extrapolations of wire assimilation times for cored wire used in the lab-scale experiments, as a function of bath temperature, convective heat transfer coefficient, and magnitude of contact thermal resistances.</p>
Full article ">
13 pages, 6149 KiB  
Article
The Integrated Preparation of Porous Tungsten Gradient Materials with a Wide Porosity Range
by Ke Zhu, Hao Jia, Jin Huang and Jian Zhang
Metals 2024, 14(4), 427; https://doi.org/10.3390/met14040427 - 5 Apr 2024
Cited by 2 | Viewed by 1398
Abstract
Porous tungsten gradient materials with ordered gradient variations in pore size have significant application value in the field of vacuum electronic devices. This work combines tape casting and dealloying methods to achieve the integrated preparation of porous tungsten gradient materials with a wide [...] Read more.
Porous tungsten gradient materials with ordered gradient variations in pore size have significant application value in the field of vacuum electronic devices. This work combines tape casting and dealloying methods to achieve the integrated preparation of porous tungsten gradient materials with a wide range of controllable porosity. The study focused on the phase composition and microstructure evolution during the preparation of porous tungsten gradient materials. The results show that the tape casting process allows for the precise and controllable thickness of each layer of the porous tungsten materials and uniform composition structure, while the stepwise dealloying of Fe and Ti enables a wide range of controllable porosity for the porous tungsten gradient materials. PVB, after thermal decomposition, provides a carbon source for the in situ reaction to form W-Fe-C compounds, and the surface diffusion behavior of W-Fe-C compounds at high temperatures improves the stratification of the porous tungsten gradient materials. This work provides a design concept for the integrated preparation of porous metal gradient materials. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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Figure 1

Figure 1
<p>Flow chart for the preparation of gradient porous tungsten: (<b>a</b>) preparation of gradient blanks and (<b>b</b>) sintering and dealloying.</p>
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<p>SEM images of W-Ti-Fe multiphase tapes with different components: (<b>a</b>) 80 at.% Fe; (<b>b</b>) 70 at.% Fe; (<b>c</b>) 60 at.% Fe; (<b>d</b>) 50 at.% Fe; (<b>e</b>) 40 at.% Fe; (<b>f</b>) 30 at.% Fe; and the atomic ratio of W to Ti is 3:1.</p>
Full article ">Figure 3
<p>TG-DSC curves of W-Ti-Fe multiphase tape.</p>
Full article ">Figure 4
<p>SEM images of W-Ti-Fe mixed powder (<b>a</b>); EDS elemental maps of W (<b>b</b>); Fe (<b>c</b>); Ti (<b>d</b>); C (<b>e</b>); and TEM image of W-Ti-Fe mixed powder (<b>f</b>).</p>
Full article ">Figure 5
<p>(<b>a</b>) FT-IR patterns of W-Ti-Fe cast tapes and mixed powder after pyrolysis; (<b>b</b>) XRD patterns of W-Ti-Fe mixed powder after pyrolysis.</p>
Full article ">Figure 6
<p>SEM images of the polished surface and the cross-section of W-Ti-Fe precursor sintered at different temperatures: (<b>a</b>,<b>d</b>) 700 °C, (<b>b</b>,<b>e</b>) 750 °C, and (<b>c</b>,<b>f</b>) 800 °C.</p>
Full article ">Figure 7
<p>SEM images of (<b>a</b>) W-Ti-Fe precursor, (<b>b</b>) prefabricated porous framework, and (<b>c</b>) W-Ti composite framework with gradient Fe content or porosity distribution. The white arrows indicate the interface between the different layers.</p>
Full article ">Figure 8
<p>(<b>a</b>) TEM image of the W-Ti composite framework; (<b>b</b>–<b>d</b>) are the corresponding EDS images of the W, Ti, and Fe elements.</p>
Full article ">Figure 9
<p>SEM images of porous tungsten porosity gradient material: (<b>a</b>) overall structure; (<b>b</b>–<b>e</b>) the enlarged images corresponding to different regions in (<b>a</b>).</p>
Full article ">Figure 10
<p>Porosity of monolayer porous tungsten with different Fe contents.</p>
Full article ">
14 pages, 7073 KiB  
Article
The On-Line Identification and Location of Welding Interference Based on CEEMD
by Peng Yu, Haichao Song, Yukuo Tian, Juan Dong, Guocheng Xu, Mingming Zhao and Xiaopeng Gu
Metals 2024, 14(4), 396; https://doi.org/10.3390/met14040396 - 28 Mar 2024
Viewed by 963
Abstract
The welding process itself is a non-linear, multivariable, coupled physical metallurgical process that is easily perturbed. Improper welding parameter selection and welding process conditions will interfere with the welding process and affect the final welding quality. This study aims to identify and locate [...] Read more.
The welding process itself is a non-linear, multivariable, coupled physical metallurgical process that is easily perturbed. Improper welding parameter selection and welding process conditions will interfere with the welding process and affect the final welding quality. This study aims to identify and locate two types of welding interference, insufficient shielding gas and unremoved oxidation film on the base metal surface, during the Pulse Multi-Control Gas Metal Arc Welding (PMC GMAW) process of aluminum alloy. The Characteristic Intrinsic Mode Function (IMF), which is closely related to the short circuit transition process, was obtained by applying the Complementary Ensemble Empirical Mode Decomposition (CEEMD) method to the welding current signal measured during the welding process. Time and frequency domain analysis of the acquired characteristic IMF was then performed. The experimental results demonstrated that for a stable welding process, the frequency of the characteristic IMF is concentrated within a narrow range. The frequency spectrum of the characteristic IMF exhibits distinct variations under different types of welding interference. Based on this, the chronological arrangement of characteristic IMF components’ frequency spectrum allows for locating welding interferences by analyzing their abnormal signals within the reconstructed signal sequence. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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Figure 1

Figure 1
<p>The welding current waveform in PMC GMAW.</p>
Full article ">Figure 2
<p>The flow chart of CEEMD.</p>
Full article ">Figure 3
<p>Welding current signal of trial 1.</p>
Full article ">Figure 4
<p>First six IMF components and corresponding frequency spectra of the welding current signal collected in trial 1.</p>
Full article ">Figure 4 Cont.
<p>First six IMF components and corresponding frequency spectra of the welding current signal collected in trial 1.</p>
Full article ">Figure 4 Cont.
<p>First six IMF components and corresponding frequency spectra of the welding current signal collected in trial 1.</p>
Full article ">Figure 5
<p>The frequency spectra of the characteristic IMF component of trial 2.</p>
Full article ">Figure 6
<p>The frequency spectra of the characteristic IMF component of trial 4.</p>
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<p>The frequency spectra of the characteristic IMF components of trial 1, (<b>a</b>) weld appearance. (<b>b</b>) frequency spectrum arranged in time order.</p>
Full article ">Figure 8
<p>The frequency spectra of the characteristic IMF components of trial 3. (<b>a</b>) weld appearance. (<b>b</b>) frequency spectrum arranged in time order.</p>
Full article ">Figure 9
<p>The frequency spectra of the characteristic IMF components of trial 5. (<b>a</b>) weld appearance. (<b>b</b>) frequency spectrum arranged in time order.</p>
Full article ">
27 pages, 8818 KiB  
Review
Research Progress on Thermal Conductivity of High-Pressure Die-Cast Aluminum Alloys
by Yixian Liu and Shoumei Xiong
Metals 2024, 14(4), 370; https://doi.org/10.3390/met14040370 - 22 Mar 2024
Cited by 9 | Viewed by 4012
Abstract
High-pressure die casting (HPDC) has been extensively used to manufacture aluminum alloy heat dissipation components in the fields of vehicles, electronics, and communication. With the increasing demand for HPDC heat dissipation components, the thermal conductivity of die-cast aluminum alloys is paid more attention. [...] Read more.
High-pressure die casting (HPDC) has been extensively used to manufacture aluminum alloy heat dissipation components in the fields of vehicles, electronics, and communication. With the increasing demand for HPDC heat dissipation components, the thermal conductivity of die-cast aluminum alloys is paid more attention. In this paper, a comprehensive review of the research progress on the thermal conductivity of HPDC aluminum alloys is provided. First of all, we introduce the general heat transport mechanism in aluminum alloys, including electrical transport and phonon transport. Secondly, we summarize several common die-cast aluminum alloy systems utilized for heat dissipation components, such as an Al–Si alloy system and silicon-free aluminum alloy systems, along with the corresponding composition optimizations for these alloy systems. Thirdly, the effect of processing parameters, which are significant for the HPDC process, on the thermal conductivity of HPDC aluminum alloys is discussed. Moreover, some heat treatment strategies for enhancing the thermal conductivity of die-cast aluminum alloys are briefly discussed. Apart from experimental findings, a range of theoretical models used to calculate the thermal conductivity of die-cast aluminum alloys are also summarized. This review aims to guide the development of new high-thermal-conductivity die-cast aluminum alloys. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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Figure 1

Figure 1
<p>Density and thermal conductivity of several pure metals, adapted from [<a href="#B9-metals-14-00370" class="html-bibr">9</a>,<a href="#B10-metals-14-00370" class="html-bibr">10</a>,<a href="#B11-metals-14-00370" class="html-bibr">11</a>].</p>
Full article ">Figure 2
<p>The process of HPDC: (<b>a</b>) pouring the melt; (<b>b</b>) slow-shot filling; (<b>c</b>) melt at the gate; (<b>d</b>) fast-shot filling; (<b>e</b>) pressure intensification; (<b>f</b>) opening the mold (reprinted with permission from ref. [<a href="#B16-metals-14-00370" class="html-bibr">16</a>], Xiaobo Li, Tsinghua Univeristy, 2017).</p>
Full article ">Figure 3
<p>Applications of HPDC heat dissipation components.</p>
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<p>The solidification rate of die casting (reprinted with permission from ref. [<a href="#B22-metals-14-00370" class="html-bibr">22</a>], 2022, Elsevier).</p>
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<p>Electron scattering patterns in an aluminum alloy.</p>
Full article ">Figure 6
<p>The effect of alloying elements on (<b>a</b>) electrical conductivity (reprinted with permission from ref. [<a href="#B62-metals-14-00370" class="html-bibr">62</a>], 2006, Elsevier) and thermal conductivity using (<b>b</b>) theoretical method (reprinted with permission from ref. [<a href="#B58-metals-14-00370" class="html-bibr">58</a>], 2023, Springer) and (<b>c</b>) experimental method of aluminum alloys (reprinted with permission from ref. [<a href="#B64-metals-14-00370" class="html-bibr">64</a>], 2015, Springer).</p>
Full article ">Figure 7
<p>Problems of traditional Al–Si die-cast alloys.</p>
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<p>Microstructure and properties of different Al–Si die-cast alloys with low Si contents: (<b>a</b>) Al–8Si (reprinted with permission from ref. [<a href="#B74-metals-14-00370" class="html-bibr">74</a>], 2018, Elsevier), (<b>b</b>) Al–6Si with Cu and Zn (reprinted with permission from ref. [<a href="#B75-metals-14-00370" class="html-bibr">75</a>], 2016, Springer), (<b>c</b>) Al–(6~8)Si with Cu or Mg (reprinted with permission from ref. [<a href="#B80-metals-14-00370" class="html-bibr">80</a>], 2022, Elsevier).</p>
Full article ">Figure 9
<p>The ternary eutectic point of the Al–Si–Ni system (reprinted with permission from ref. [<a href="#B84-metals-14-00370" class="html-bibr">84</a>], 2015, Springer).</p>
Full article ">Figure 10
<p>(<b>a</b>) Effect of Mn on thermal conductivity of the die-cast Al–Si alloy (reprinted with permission from ref. [<a href="#B61-metals-14-00370" class="html-bibr">61</a>], 2013, Springer). (<b>b</b>) The purification process of boron treatment (reprinted with permission from ref. [<a href="#B91-metals-14-00370" class="html-bibr">91</a>], 2018, Elsevier) and (<b>c</b>) the effect of boron on the thermal conductivity of ADC12 alloy (reprinted ref. [<a href="#B95-metals-14-00370" class="html-bibr">95</a>]).</p>
Full article ">Figure 11
<p>(<b>a</b>,<b>b</b>) Mechanism of improving thermal conductivity by modifying the eutectic Si (reprinted with permission from ref. [<a href="#B24-metals-14-00370" class="html-bibr">24</a>], 2020, Springer).</p>
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<p>(<b>a</b>) Variation in thermal conductivity under different fractions of porosity caused by processing parameters (reprinted with permission from ref. [<a href="#B123-metals-14-00370" class="html-bibr">123</a>], 2017, Elsevier). (<b>b</b>) Thermal conductivity and corresponding (<b>c</b>) porosity distribution under different vacuum levels (reprinted with permission from ref. [<a href="#B124-metals-14-00370" class="html-bibr">124</a>], 2020, Elsevier). (<b>d</b>) Distribution of ESCs and (<b>e</b>) thermal conductivity of die-cast alloys under different shot sleeves (reprinted with permission from ref. [<a href="#B125-metals-14-00370" class="html-bibr">125</a>], 2022, Elsevier).</p>
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<p>The schematic diagrams of the ACSR Rheo-HPDC process (reprinted with permission from ref. [<a href="#B36-metals-14-00370" class="html-bibr">36</a>], 2022, Elsevier).</p>
Full article ">
16 pages, 3834 KiB  
Article
A Fundamental Study on the Preparation of Sodium Tungstate from Wolframite via the Smelting Process
by Liqiang Xu and Baojun Zhao
Metals 2024, 14(3), 299; https://doi.org/10.3390/met14030299 - 1 Mar 2024
Cited by 1 | Viewed by 1633
Abstract
Tungsten is a high-value resource with a wide range of applications. The tungsten metal is produced via ammonium paratungstate, which is a multi-stage process including leaching, conversion, precipitation, calcination, and reduction. A short process to produce tungsten metal from the electrolysis of molten [...] Read more.
Tungsten is a high-value resource with a wide range of applications. The tungsten metal is produced via ammonium paratungstate, which is a multi-stage process including leaching, conversion, precipitation, calcination, and reduction. A short process to produce tungsten metal from the electrolysis of molten sodium tungstate has been demonstrated. However, sodium tungstate cannot be directly produced from wolframite in the conventional hydrometallurgical process. There was no information reported in the literature on producing sodium tungstate directly from tungsten concentrates. The present study proposed a simple and low-cost process to produce sodium tungstate by high-temperature processing of wolframite. The mixtures of wolframite, sodium carbonate, and silica were melted in air between 1100 and 1300 °C. High-density sodium tungstate was easily separated from the immiscible slag, which contained all impurities from wolframite, flux, excess sodium oxide, and dissolved tungsten oxide. The slag was further water leached to recover sodium tungstate in the solution. Effects of Na2CO3/Ore and SiO2/Ore ratios, temperature, and reaction time on the recovery of tungstate and the purity of sodium tungstate were systematically studied. Sodium tungstate containing over 78% WO3 was produced in the smelting process, which is suitable for the electrolysis process. The experimental results will provide a theoretical basis for the direct production of sodium tungstate from wolframite. The compositions of the WO3-containing slags and sodium tungstate reported in the present study fill the knowledge gap of the tungsten-containing thermodynamic database. Further studies to use complex and low-grade tungsten concentrates to produce sodium tungstate are underway. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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Figure 1

Figure 1
<p>Experimental procedure in the present study.</p>
Full article ">Figure 2
<p>Effect of Na<sub>2</sub>CO<sub>3</sub>/Ore on (<b>a</b>) direct recovery of WO<sub>3</sub> and (<b>b</b>) WO<sub>3</sub> and SiO<sub>2</sub> contents in sodium tungstate, SiO<sub>2</sub>/Ore = 0.3, 1200 °C, 60 min.</p>
Full article ">Figure 3
<p>Effect of Na<sub>2</sub>CO<sub>3</sub>/Ore on total recovery of WO<sub>3</sub>, SiO<sub>2</sub>/Ore = 0.3, 1200 °C, 60 min.</p>
Full article ">Figure 4
<p>Effect of Na<sub>2</sub>CO<sub>3</sub>/Ore on direct recovery and total recovery of WO<sub>3</sub>, SiO<sub>2</sub>/Ore = 0.3, 1200 °C, 60 min.</p>
Full article ">Figure 5
<p>Effect of SiO<sub>2</sub>/Ore on (<b>a</b>) direct recovery of WO<sub>3</sub> and (<b>b</b>) WO<sub>3</sub> and SiO<sub>2</sub> contents in sodium tungstate, Na<sub>2</sub>CO<sub>3</sub>/Ore = 1.1, 1200 °C, 60 min.</p>
Full article ">Figure 6
<p>Effect of SiO<sub>2</sub>/Ore on total recovery of WO<sub>3</sub>, Na<sub>2</sub>CO<sub>3</sub>/Ore = 1.1, 1200 °C, 60 min.</p>
Full article ">Figure 7
<p>Effect of SiO<sub>2</sub>/Ore on direct recovery and total recovery of WO<sub>3</sub>, Na<sub>2</sub>CO<sub>3</sub>/Ore = 1.1, 1200 °C, 60 min.</p>
Full article ">Figure 8
<p>Effect of temperature on (<b>a</b>) direct recovery of WO<sub>3</sub> and (<b>b</b>) WO<sub>3</sub> and SiO<sub>2</sub> contents in sodium tungstate, Na<sub>2</sub>CO<sub>3</sub>/Ore = 1.1, SiO<sub>2</sub>/Ore = 0.3, 60 min.</p>
Full article ">Figure 9
<p>Effect of temperature on total recovery of WO<sub>3</sub>, Na<sub>2</sub>CO<sub>3</sub>/Ore = 1.1, SiO<sub>2</sub>/Ore = 0.3, 60 min.</p>
Full article ">Figure 10
<p>Effect of temperature on direct recovery and total recovery of WO<sub>3</sub>, Na<sub>2</sub>CO<sub>3</sub>/Ore = 1.1, SiO<sub>2</sub>/Ore = 0.3, 60 min.</p>
Full article ">Figure 11
<p>Effect of reaction time on (<b>a</b>) direct recovery of WO<sub>3</sub> and (<b>b</b>) WO<sub>3</sub> and SiO<sub>2</sub> contents in sodium tungstate, Na<sub>2</sub>CO<sub>3</sub>/Ore = 1.1, SiO<sub>2</sub>/Ore = 0.3, 1200 °C.</p>
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<p>Effect of reaction time on total recovery of WO<sub>3</sub>, Na<sub>2</sub>CO<sub>3</sub>/Ore = 1.1, SiO<sub>2</sub>/Ore = 0.3, 1200 °C.</p>
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<p>Effect of reaction time on direct recovery and total recovery of WO<sub>3</sub> and WO<sub>3</sub> content in sodium tungstate, Na<sub>2</sub>CO<sub>3</sub>/Ore = 1.1, SiO<sub>2</sub>/Ore = 0.3, 1200 °C.</p>
Full article ">
14 pages, 2723 KiB  
Article
Comparative Study of Anti-Corrosion Properties of Different Types of Press-Hardened Steels
by Hao Peng, Yunlong Zhao, Wanwan Fu, Zhishan Chen, Man Zhang, Jiesheng Liu and Xiaoming Tan
Materials 2024, 17(5), 1022; https://doi.org/10.3390/ma17051022 - 23 Feb 2024
Cited by 1 | Viewed by 1210
Abstract
Hot stamping (or press hardening) is a new technology that is widely used in the production of advanced high-strength steel parts for automotive applications. Electrochemical measurements, including potentiodynamic polarization and electrochemical impedance spectroscopy (EIS), and accelerated corrosion tests (the neutral salt spray test [...] Read more.
Hot stamping (or press hardening) is a new technology that is widely used in the production of advanced high-strength steel parts for automotive applications. Electrochemical measurements, including potentiodynamic polarization and electrochemical impedance spectroscopy (EIS), and accelerated corrosion tests (the neutral salt spray test and periodic immersion test) were conducted on press-hardened samples produced from uncoated (cold-rolled and cold strip production (CSP) hot-rolled) and Al–Si-coated press-hardened steels to elucidate their distinct anti-corrosion mechanisms. The cross-sectional micromorphology and element distribution of three types of press-hardened steels after a neutral salt spray test were observed using scanning electron microscopy (SEM) and energy-dispersive X-ray analysis (EDAX). The corrosion resistance of Al–Si-coated press-hardened steel was found to be significantly diminished following the hot stamping process due to the presence of microcracks and elevated iron content in the coating subsequent to austenitizing heat treatment. On the other hand, the corrosion resistance of uncoated press-hardened sheets produced from cold-rolled and CSP hot-rolled press-hardened steel was found to be proximal due to their nearly identical composition and microstructure (fully martensite) after the hot stamping process. Considering the high efficiency and energy-saving properties of hot-rolled press-hardened steel, it holds the potential to replace cold-rolled and even aluminum–silicon-coated press-hardened steel in automobile manufacturing. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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Figure 1
<p>Potentiodynamic polarization curves of press-hardened steel samples (cold-rolled, CSP hot-rolled, and Al–Si-coated).</p>
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<p>Nyquist plots of press-hardened sheets made from cold-rolled and CSP hot-rolled press-hardened steels in a 3.5 wt.% NaCl solution.</p>
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<p>Surface appearances of unpainted press-hardened samples following the NSST for 9 h [<a href="#B9-materials-17-01022" class="html-bibr">9</a>].</p>
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<p>Surface appearances of the Al–Si-coated sample (after exposure to salt spray for a certain amount of time) before and after hot stamping process.</p>
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<p>Comparison of Al–Si-layered steel before and after hot stamping process.</p>
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<p>Cross-sectional microstructure and EDAX charts of the press-hardened Al–Si-coated sample. (<b>a</b>) Sample point A; (<b>b</b>) Sample point B; (<b>c</b>) Sample point C.</p>
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<p>Cross-sectional microstructures of three kinds of press-hardened samples following 9 h of NSST exposure.</p>
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<p>Corrosion depths of three kinds of press-hardened samples following the NSST for 9 h.</p>
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10 pages, 24169 KiB  
Article
Effects of Zn Addition and Twin Roll Casting Process on the Microstructure, Texture, and Mechanical Properties of the Mg-Al-Mn-Ca Sheet
by Donghwan Eom, Sangbong Yi, Dietmar Letzig and No-Jin Park
Metals 2024, 14(3), 261; https://doi.org/10.3390/met14030261 - 22 Feb 2024
Cited by 1 | Viewed by 1350
Abstract
In this work, the microstructure and texture of Mg-1.0Al-xZn-0.2Mn-0.5Ca (wt.%, x = 0, 1) alloys, which were produced via conventional casting or twin roll casting (TRC), were investigated, and their relation to the mechanical properties of the sheets at the final gage was [...] Read more.
In this work, the microstructure and texture of Mg-1.0Al-xZn-0.2Mn-0.5Ca (wt.%, x = 0, 1) alloys, which were produced via conventional casting or twin roll casting (TRC), were investigated, and their relation to the mechanical properties of the sheets at the final gage was analyzed. In the Zn-containing AZMX1100 alloy sheets, the amount and size of the secondary phases were significantly reduced, in comparison to the Zn-free AMX100 alloy sheet. The TRC sheet shows a smaller grain structure and fine secondary phases in comparison to the sheets produced via the conventional casting process. The texture of the AMX100 sheet is characterized by the basal poles tilted in the sheet rolling direction (RD). In the AZMX1100 sheets, the texture with the tilted basal poles towards the RD and transverse direction (TD) was developed after recrystallization annealing, while the tilting angle of the basal pole in the TD is larger than in the RD. There is no significant difference in the texture between the sheets produced by the casting and TRC process. The highest yield strength was obtained in the AZMX1100 sheet produced by the TRC process, and all examined sheets showed the mechanical anisotropy in accordance with their textures. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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<p>Optical micrographs of the cast blocks of (<b>a</b>) AMX100 and (<b>b</b>) AZMX1100 alloy.</p>
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<p>Optical micrographs of the as-TRC strip of the AZMX1100 alloy with a thickness of 4.8 mm at the (<b>a</b>) ND-RD plane, (<b>b</b>) centerline segregation, and (<b>c</b>) columnar structure inclined in the RD at the near-surface area. The sample directions and planes are schematically shown.</p>
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<p>SEM images of the as-TRC strip of the AZMX1100 alloy at the (<b>a</b>) ND-RD plane, (<b>b</b>) secondary phase particle formed at the boundary of columnar grains, (<b>c</b>) secondary phase particle formed at the centerline segregation. EDS results analyzed on the rectangular area marked on (<b>a</b>), and the secondary phases marked on (<b>b</b>,<b>c</b>) are listed in the tables.</p>
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<p>Optical micrographs of the hot-rolled sheets of (<b>a</b>) casting AMX100, (<b>b</b>) casting AZMX1100, (<b>c</b>) TRC AZMX1100 alloys, and after the heat treatment at 350 °C for 2 h of (<b>d</b>) casting AMX100, (<b>e</b>) casting AZMX1100, and (<b>f</b>) TRC AZMX1100 sheets. The arrows indicate the secondary phases.</p>
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<p>Grain size distribution graph of the as-rolled sheets of (<b>a</b>) casting AMX100, (<b>b</b>) casting AZMX1100, (<b>c</b>) TRC AZMX1100, and after the heat treatment at 350 °C for 2 h of (<b>d</b>) casting AMX100, (<b>e</b>) casting AZMX1100, and (<b>f</b>) TRC AZMX1100 sheets.</p>
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<p>SEM images of the hot-rolled sheets of (<b>a</b>) casting AMX100, (<b>b</b>) casting AZMX1100, (<b>c</b>) TRC AZMX1100 alloys. EDS results analyzed on the secondary phases found in the rolled sheets are listed in the table. EDS area mapping data of the hot-rolled sheets of (<b>d</b>) casting AZMX1100. (Different magnifications were employed due to the large differences in size of the secondary phases between the sheets).</p>
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<p>Phase diagram of Mg-1.0Al-(0~1)Zn-0.2Mn-0.5Ca, in wt.%, calculated using thermodynamic simulation software Pandat (<a href="https://computherm.com/software" target="_blank">https://computherm.com/software</a>, accessed on 17 February 2024) with PanMg_TH (CompuTherm LLC, Middleton, MI, USA).</p>
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<p>Recalculated (0002) pole figures of the hot-rolled sheets of (<b>a</b>) casting AMX100, (<b>b</b>) casting AZMX1100, (<b>c</b>) TRC AZMX1100 alloys, and the rolled sheet after the heat treatment at 350 °C for 2 h at (<b>d</b>) casting AMX100, (<b>e</b>) casting AZMX1100, and (<b>f</b>) TRC AZMX1100 (Maximum and minimum pole density in multiples of random distribution, m.r.d., X: RD, Y: TD).</p>
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<p>Stress–strain curves of the sheets after the heat treatment at 350 °C for 2 h.</p>
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14 pages, 8415 KiB  
Article
Dissolution of CaO in SiO2-CaO-Al2O3 Slag in Si Production
by Marit Buhaug Folstad, Kristian Etienne Einarsrud and Merete Tangstad
Metals 2024, 14(2), 243; https://doi.org/10.3390/met14020243 - 16 Feb 2024
Viewed by 1572
Abstract
This work investigates the dissolution of CaO into three different compositions of SiO2-CaO-Al2O3 slag similar to those found in industrial Si furnaces. It was found that CaO dissolution into the slag is fast. During the dissolution process, a [...] Read more.
This work investigates the dissolution of CaO into three different compositions of SiO2-CaO-Al2O3 slag similar to those found in industrial Si furnaces. It was found that CaO dissolution into the slag is fast. During the dissolution process, a layer containing 35–42% CaO was formed between the CaO particle and the slag, which corresponded to the phases CaO·Al2O3·2SiO2 or 2CaO·Al2O3·SiO2 in this study. Two models were investigated to determine the dissolution rate of the three slags. In the first model, the CaO particle is assumed to be a smooth shrinking sphere, and the rate controlled by the chemical reaction rate. The second model assumes that the rate is controlled by mass transport and depends on the diffusion rate of CaO through a boundary layer on the surface of the CaO. Both models gave similar accuracy to the experimental values, and a proportional relationship between the rate constants and the viscosities was obtained. At 1500 °C, the diffusion coefficients were found to be in the order of 10−6 cm2/s. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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<p>Illustration of a Si furnace including the main zones and approximate temperatures at different positions. The temperatures are based on modeled temperatures from Myrhaug [<a href="#B2-metals-14-00243" class="html-bibr">2</a>].</p>
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<p>Heating profiles for the dissolution experiments in the graphite tube furnace shown as the blue, green and grey lines. The red line displays the heating profile for the sessile drop experiments at 1500 °C. The circles represent the different holding times, and “*2” and “*3” means that the experiment is repeated 2 and 3 times, respectively.</p>
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<p>Schematic overview of the sessile drop furnace.</p>
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<p>Slag and CaO before sessile drop experiment and slag after dissolution experiment.</p>
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<p>Schematic of the observed dissolution of a CaO particle into 56% SiO<sub>2</sub>—15% CaO—29% Al<sub>2</sub>O<sub>3</sub> at 1600 °C for different holding times in the graphite tube furnace.</p>
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<p>CT images from graphite tube furnace experiment of the second parallel heated to 1600 °C with heating rate 50 °C/min up to 1200 °C and 25 °C/min up to 1600 °C with 0 min holding time. To the left is an image from the top of the sample, and the right image is a vertical cross section in the center of the sample. The yellow line in the vertical section marks the height of the horizontal image. More CaO in the slag can be observed as the brighter grey color. The darkest grey color is the graphite crucible.</p>
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<p>Results of WDS analysis vertical section for the samples with heating rate 50 °C/min to 1200 °C and 25 °C/min to 1600 °C. The CaO/slag ratio was 0.1.</p>
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<p>Cross section images and EPMA images after experiments for the 1500 °C experiments with CaO and slag 1 with 0, 5 and 10 min holding time.</p>
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<p>Dissolution curves for (<b>a</b>) slag 1, (<b>b</b>) slag 2 and (<b>c</b>) slag 3 for the experiments at 1500 °C with 0, 5 and 10 min holding time.</p>
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<p>Modeled dissolution rate curves for (<b>a</b>) slag 1, (<b>b</b>) slag 2 and (<b>c</b>) slag 3 using shrinking sphere model and mass transport model. The curves are compared together with the experimental results.</p>
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<p>Viscosity with increasing %CaO for slag 1, 2 and 3 at temperature 1600 °C and 1800 °C. The viscosities are calculated using FactSage 8.1.</p>
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15 pages, 1902 KiB  
Article
Pilot Tests of Pre-Reduction in Chromium Raw Materials from Donskoy Ore Mining and Processing Plant and Melting of High-Carbon Ferrochromium
by Yerbol Shabanov, Yerbolat Makhambetov, Zhalgas Saulebek, Ruslan Toleukadyr, Sailaubai Baisanov, Nurzhan Nurgali, Azamat Shotanov, Murat Dossekenov and Yerlan Zhumagaliyev
Metals 2024, 14(2), 202; https://doi.org/10.3390/met14020202 - 6 Feb 2024
Cited by 4 | Viewed by 1687
Abstract
Experiments were conducted to pilot the initial reduction in chromium raw materials using the innovative Hoganas technology in a tunnel furnace. To simulate the process, a gas-fired bogie hearth furnace was employed. Technological containers made of silicon carbide crucibles were utilized. Sixteen different [...] Read more.
Experiments were conducted to pilot the initial reduction in chromium raw materials using the innovative Hoganas technology in a tunnel furnace. To simulate the process, a gas-fired bogie hearth furnace was employed. Technological containers made of silicon carbide crucibles were utilized. Sixteen different combinations of ore and coal mixtures were employed for the initial reduction process. Their total mass was more than 20 tons. Their heat treatment was performed at different temperatures and durations. During the pilot tests, the possibility of achieving chromium metallization was confirmed. Thus, it explains the application of a pre-reduction instead of the sintering or charge heating before the ferrochromium melting, i.e., the power consumption is minimized during the final remelting of the product in DC furnaces. The pilot melting of three batches of the pre-reduced chromium raw materials with various chromium metallization degrees has been tested in the ore-smelting furnace at Zh. Abishev Chemical–Metallurgical Institute (Karaganda). The capacity was 0.2 MVA. To evaluate the technical and economic efficiency of remelting pre-reduced chromium raw materials in commercial DC furnaces, a specific batch of primary ingredients for producing high-carbon ferrochromium, including chromite ore, coke, and quartz flux, was successfully melted in a segregated phase. As a result of the study, it was found that the specific energy consumption for melting high-carbon ferrochromium in the pilot furnace depends on the chromium metallization degree. The researchers tested a range of chromium metallization degrees from 0 to 65% and determined the corresponding specific energy consumption for each degree. Using the data obtained from the study, the researchers were able to assess the melting indexes of high-carbon ferrochromium in a larger 72 MW furnace. They found that by using a pre-reduced product with a chromium metallization degree of 65%, it was possible to reduce the specific energy consumption by half, to around 3.4 MW·h per ton of chromium. Overall, this study highlights the importance of considering the chromium metallization degree when determining the specific energy consumption for melting high-carbon ferrochromium. By optimizing the metallization degree, significant energy savings can be achieved, leading to more efficient and sustainable production processes. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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<p>The total energy consumption for melting high-carbon ferrochromium through pre-reduction is influenced by the level of chromium pre-metallization in both the ore-smelting furnace and DC furnace. Reprinted with permission from ref. [<a href="#B14-metals-14-00202" class="html-bibr">14</a>]. 2022 Springer Nature.</p>
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<p>The calculated ratio between the metallization of Fe and Cr during the preliminary reduction in Kazakhstani chromites. Reprinted with permission from ref. [<a href="#B14-metals-14-00202" class="html-bibr">14</a>]. 2022 Springer Nature.</p>
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<p>Ore smelting furnace with a capacity of 200 kVA: (<b>a</b>) view of the furnace top and (<b>b</b>) release of metal into a cascade of molds. Reprinted with permission from ref. [<a href="#B15-metals-14-00202" class="html-bibr">15</a>]. 2023 Springer Nature.</p>
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<p>Products of smelting high-carbon ferrochrome in a pilot furnace: (<b>a</b>) metal and slag ingots in a mold and (<b>b</b>) metal after cooling.</p>
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<p>Dependence of specific energy consumption on the metallization degree of pre-reduced product during melting of high-carbon ferrochromium.</p>
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13 pages, 1270 KiB  
Article
Thermodynamic Modeling of the Drowning-Out Crystallization Process for LiOH and CHLiO2
by Raquel González, Yahaira Barrueto and Yecid P. Jiménez
Metals 2024, 14(1), 78; https://doi.org/10.3390/met14010078 - 9 Jan 2024
Cited by 1 | Viewed by 1789
Abstract
This study focuses on the thermodynamic modeling of the crystallization by the drowning process for two lithium salts: lithium hydroxide (LiOH) and lithium formate (CHLiO2). The modeling involves utilizing thermodynamic properties, such as the activity, osmotic, and solubility coefficients, within the [...] Read more.
This study focuses on the thermodynamic modeling of the crystallization by the drowning process for two lithium salts: lithium hydroxide (LiOH) and lithium formate (CHLiO2). The modeling involves utilizing thermodynamic properties, such as the activity, osmotic, and solubility coefficients, within the ternary systems of LiOH + cosolvent + water and CHLiO2 + cosolvent + water, as well as their respective binary constituent systems. Ethanol is chosen as the cosolvent for both salts, facilitating a comparative analysis. Given the limited availability of thermodynamic data for lithium formate with different cosolvents, the study aims to address this gap. The modified Pitzer model was employed for the modeling process, where the parameters were successfully obtained for both systems, with a deviation of less than 1%. Additionally, the mass and energy balance for the drowning-out crystallization process of both salts was performed. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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<p>Comparison of osmotic coefficient data for the LiOH + H<sub>2</sub>O system at 298.15 K. [<a href="#B1-metals-14-00078" class="html-bibr">1</a>] –●–, [<a href="#B27-metals-14-00078" class="html-bibr">27</a>] –♦–, and [<a href="#B28-metals-14-00078" class="html-bibr">28</a>] –■–.</p>
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<p>Comparison of activity coefficient data for the LiOH + H<sub>2</sub>O system at 298.15 K. [<a href="#B1-metals-14-00078" class="html-bibr">1</a>] –●–, [<a href="#B27-metals-14-00078" class="html-bibr">27</a>] –♦–, and [<a href="#B28-metals-14-00078" class="html-bibr">28</a>] –■–.</p>
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<p>Lithium formate solubility data for the ternary system: 313.15 K ▬, 303.15 K ■, 298.15 K ♦, 293.15 K ▲, and 283.15 K ●. Data from [<a href="#B29-metals-14-00078" class="html-bibr">29</a>].</p>
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<p>Comparison of lithium formate precipitate results at different temperatures: 283.15 K ●, 293.15 K ■, 298.15 K ♦, 303.15 K ▲, and 313.15 K ▬. Curves are guides for the eye.</p>
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17 pages, 3401 KiB  
Article
Two-Stage Leaching of PCBs Using Sulfuric and Nitric Acid with the Addition of Hydrogen Peroxide and Ozone
by Magdalena Lisińska, Tomasz Wojtal, Mariola Saternus, Joanna Willner, Martyna Rzelewska-Piekut and Krzysztof Nowacki
Materials 2024, 17(1), 219; https://doi.org/10.3390/ma17010219 - 30 Dec 2023
Cited by 4 | Viewed by 2127
Abstract
The paper presents the possibility of recovering metals from printed circuit boards (PCBs) of spent mobile phones using the hydrometallurgical method. Two-stage leaching of Cu(II), Fe(III), Sn(IV), Zn(II), Ni(II) and Pb(II) with H2SO4 (2 and 5 M) and HNO3 [...] Read more.
The paper presents the possibility of recovering metals from printed circuit boards (PCBs) of spent mobile phones using the hydrometallurgical method. Two-stage leaching of Cu(II), Fe(III), Sn(IV), Zn(II), Ni(II) and Pb(II) with H2SO4 (2 and 5 M) and HNO3 (2 M) with the addition of H2O2 (10 and 30%) and O3 (9 or 15 g/h) was conducted at various process conditions (temperature—313, 333 and 353 K, time—60, 120, 240, 300 min, type and concentration of leaching agent, type and concentration of oxidant, solid–liquid ratio (S/L)), allowing for a high or total metals leaching rate. The use of two leaching stages allows for the preservation of selectivity, separation and recovery of metals: in the first stage of Fe(III), Sn(IV) and in the second stage of the remaining tested metal ions, i.e., Cu(II), Zn(II), Ni(II) and Pb(II). Removing Fe from the tested PCBs’ material at the beginning of the process eliminates the need to use magnetic methods, the purpose of which is to separate magnetic metal particles (ferrous) from non-magnetic (non-ferrous) particles; these procedures involve high operating costs. Since the leaching of Cu(II) ions with sulfuric(VI) acid practically does not occur (less than 1%), this allows for almost complete transfer of these ions into the solution in the second stage of leaching. Moreover, to speed up the process and not generate too many waste solutions, oxidants in the form of hydrogen peroxide and ozone were used. The best degree of leaching of all tested metal ions was obtained when 2 M sulfuric(VI) acid at 353 K was used in the 1st research stage, and 2 M nitric(V) acid and 9 g/h O3 at 298 K in the 2nd stage of leaching, which allowed it to be totally leached 100% of Fe(III), Cu(II), Sn(IV), Zn(II), Ni(II) and 90% Pb(II). Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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<p>Scheme of two-stage leaching of metals from PCBs with the use of: 1st stage—2 M H<sub>2</sub>SO<sub>4</sub> at 353 K, 2nd stage—2 M HNO<sub>3</sub> with the addition of H<sub>2</sub>O<sub>2</sub> and O<sub>3</sub> as the oxidants at 298 K.</p>
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<p>Influence of H<sub>2</sub>SO<sub>4</sub> concentration on the leaching efficiency of: (<b>a</b>) Fe(III), (<b>b</b>) Sn(IV), (<b>c</b>) Zn(II), (<b>d</b>) Ni(II) from PCBs (temperature: 313–353 K, time: 480 min).</p>
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<p>Dependence of the leaching efficiency of: (<b>a</b>) Fe(III), (<b>b</b>) Sn(IV), (<b>c</b>) Zn(II), (<b>d</b>) Ni(II) on temperatures of 313, 333, 353 K for 2 and 5 M H<sub>2</sub>SO<sub>4</sub> (leaching time: 480 min).</p>
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<p>Leaching efficiency of Fe(III), Sn(IV), Zn(II), Ni(II) from PCBs with 2 M H<sub>2</sub>SO<sub>4</sub> (temperature 353 K, leaching time: 300 min).</p>
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<p>Leaching efficiency as a function of time for: Fe(II) and Sn(IV) (temperature 353 K, leaching solution: 2 M H<sub>2</sub>SO<sub>4</sub>).</p>
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<p>Linear model for: Fe(II) and Sn(IV) (leaching solution: 2 M H<sub>2</sub>SO<sub>4</sub>, leaching time: 300 min).</p>
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<p>Influence of the oxidizing agent H<sub>2</sub>O<sub>2</sub> and O<sub>3</sub> on the leaching efficiency of: (<b>a</b>) Cu(II), (<b>b</b>) Pb(II) from PCBs (temperature: 313 K, time: 300 min).</p>
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<p>Degree of leaching of metal ions after two-stage leaching process ((<span style="color:#767171">■</span>)—1st stage—2 M H<sub>2</sub>SO<sub>4</sub>, (<span style="color:#2F5496">■</span>)—2nd stage—(<b>a</b>) 2 M HNO<sub>3</sub>, (<b>b</b>) 2 M HNO<sub>3</sub>+10% H<sub>2</sub>O<sub>2</sub>, (<b>c</b>) 2 M HNO<sub>3</sub>+30% H<sub>2</sub>O<sub>2</sub>, (<b>d</b>) 2 M HNO<sub>3</sub>+9 g/h O<sub>3</sub>, (<b>e</b>) 2 M HNO<sub>3</sub>+15 g/h O<sub>3</sub>).</p>
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15 pages, 3840 KiB  
Article
Mineral Magnetic Modification of Fine Iron Ore Tailings and Their Beneficiation in Alternating Magnetic Fields
by Nataliia Dudchenko, Vitalii Ponomar, Volodymyr Ovsiienko, Yurii Cherevko and Ilana Perelshtein
Metals 2024, 14(1), 26; https://doi.org/10.3390/met14010026 - 25 Dec 2023
Cited by 2 | Viewed by 1742
Abstract
In this paper, the properties, mineral magnetic modification, and beneficiation of tailings from the central mining and processing plant in Kryvyi Rih, Ukraine, have been studied. Samples were investigated by X-ray diffraction, X-ray fluorescence, microscopy, and magnetization measurements. The beneficiation was conducted using [...] Read more.
In this paper, the properties, mineral magnetic modification, and beneficiation of tailings from the central mining and processing plant in Kryvyi Rih, Ukraine, have been studied. Samples were investigated by X-ray diffraction, X-ray fluorescence, microscopy, and magnetization measurements. The beneficiation was conducted using magnetizing roasting with carbon monoxide followed by dry low-intensity magnetic separation. The effects of chemical, mineral, and granulometric composition on the processing of fine tailings of different sizes sampled at different points of the tailings pond were investigated. Additionally, we proposed a new approach for magnetic separation of fine magnetically modified tailings based on the combination of permanent and alternating magnetic fields. Magnetizing roasting resulted in an enhancement in mass magnetization to 11–62 Am2/kg in comparison with initial values of 0.3–1.5 Am2/kg. After magnetic separation, the magnetic concentrates consisted almost completely of magnetite (with the magnetization of 75–88 Am2/kg) and non-magnetic residues contained major quartz. The content of iron in magnetic concentrates reached 68.5–70.2 wt.% and iron recovery 77–96 wt.%, depending on size fraction. We could conclude that the tailings are represented by fine-grained liberated material that can be effectively upgraded using magnetizing roasting and magnetic separation into two valuable products, such as iron concentrate and quartz powder. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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<p>Scheme of a device for the separation of highly magnetic dispersed iron ores in an alternating magnetic field, where 1 is a permanent magnet; 2—an electromagnet; 3—a tripod; 4—an ore loading platform; 5—a concentrate collector; 6—the base of the separator; 7—a voltage regulator, the black circles denote magnetite, and the white circles denote quartz.</p>
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<p>Flowchart of the processing of iron ore tailings.</p>
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<p>Mineral composition and iron content of tailings materials obtained from the sampling profiles located at different distances from the slurry pipeline of the tailings pond—T1(50 m north), T3-14, and T3-15 (150 m south), T5 (350 m south), T12 (500 m south).</p>
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<p>Granulometric analysis of tailings samples T1, T5, and T12. For better data representation, the mass of initial granulometric fractions 0.05−0.63 and 0.063−0.071 mm is displayed as a sum of 0.05−0.07 mm.</p>
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<p>Mineral composition and iron content for granulometric fractions &lt;0.05 mm, 0.063–0.071 mm, and 0.1–0.25 mm of tailings samples T1, T5, and T12.</p>
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<p>The saturation magnetization of initial and roasted samples from the tailings.</p>
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<p>X-ray diffraction analysis of the original (<b>a</b>) and roasted (<b>b</b>) samples of T1 and fractions &lt;0.05 mm (T1-01), 0.063–0.071 mm (T1-03), 0.071–0.1 mm (T1-06), where <span class="html-fig-inline" id="metals-14-00026-i001"><img alt="Metals 14 00026 i001" src="/metals/metals-14-00026/article_deploy/html/images/metals-14-00026-i001.png"/></span>—quartz, <span class="html-fig-inline" id="metals-14-00026-i002"><img alt="Metals 14 00026 i002" src="/metals/metals-14-00026/article_deploy/html/images/metals-14-00026-i002.png"/></span>—hematite, and <span class="html-fig-inline" id="metals-14-00026-i003"><img alt="Metals 14 00026 i003" src="/metals/metals-14-00026/article_deploy/html/images/metals-14-00026-i003.png"/></span>—magnetite.</p>
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<p>The saturation magnetization of the roasted size fractions &lt;0.05 mm, 0.063–0.071 mm, 0.071–0.1 mm of samples T1, T5, and T12.</p>
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<p>Mass magnetization of residues and concentrates after magnetic separation of tailings material from T1 and its size fractions &lt;0.05 mm (T1-01), 0.063–0.071 mm (T1-03), 0.071–0.1 mm (T1-06).</p>
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<p>X-ray diffraction analysis of concentrates (<b>a</b>) and residues (<b>b</b>) of samples T1 and fractions &lt;0.05 mm (T1-01), 0.063–0.071 mm (T1-03), 0.071–0.1 mm (T1-06), where <span class="html-fig-inline" id="metals-14-00026-i001"><img alt="Metals 14 00026 i001" src="/metals/metals-14-00026/article_deploy/html/images/metals-14-00026-i001.png"/></span>—quartz, <span class="html-fig-inline" id="metals-14-00026-i002"><img alt="Metals 14 00026 i002" src="/metals/metals-14-00026/article_deploy/html/images/metals-14-00026-i002.png"/></span>—hematite, and <span class="html-fig-inline" id="metals-14-00026-i003"><img alt="Metals 14 00026 i003" src="/metals/metals-14-00026/article_deploy/html/images/metals-14-00026-i003.png"/></span>—magnetite.</p>
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11 pages, 5650 KiB  
Article
Structure and Heat Transfer Characteristic Evolution of CaO-SiO2-CaF2-Based Solid Mold Flux Film upon Solidification
by Xiao Long, Wenbo Luo, Xiang Li, Shaolei Long, Honggang Ma, Dayang Luo and Congxin Zheng
Metals 2024, 14(1), 1; https://doi.org/10.3390/met14010001 - 19 Dec 2023
Cited by 1 | Viewed by 1196
Abstract
In this study, two typical commercially used CaO-SiO2-CaF2-based mold fluxes with different basicities were adopted. Solid slag films of the two mold fluxes were obtained by immersing an improved water-cooled copper probe in the molten fluxes for different probe [...] Read more.
In this study, two typical commercially used CaO-SiO2-CaF2-based mold fluxes with different basicities were adopted. Solid slag films of the two mold fluxes were obtained by immersing an improved water-cooled copper probe in the molten fluxes for different probe immersion times and molten slag temperatures. The film thickness, closed porosity, and roughness of the film surfaces in contact with the copper probe were measured. The heat flux through the solidified films and the comprehensive thermal conductivity of the films were both calculated. The results indicated that compared with the heat flux through high-basicity films, the heat flux through low-basicity films exhibited high fluctuation due to the evolution of fusion cracks within the glass layer. High-basicity mold fluxes resulted in higher thickness, growth velocity, surface roughness, and devitrification velocity of the films. With the growth and crystallization of the slag films, the comprehensive thermal conductivity of the high-basicity films increased significantly. For the low-basicity films, their comprehensive thermal conductivity first decreased and then increased after the solidification time exceeded 30 s. The comprehensive thermal conductivity of the high- and low-basicity films ranged from 0.63 to 0.91 and 0.62 to 0.81 W/(m·K), respectively. The results provide a novel method for analyzing the potential effect of the structural factors of slag films on heat transfer control and controlling the heat transfer behavior of slag films. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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<p>Heat flux through slag films solidified in slag bulk with different temperatures, mold flux with basicity 1.28 (<b>a</b>) and 0.85 (<b>b</b>).</p>
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<p>Mean heat fluxes and corresponding standard deviation through slag films during solidification. Bulk slag temperature 1350 °C.</p>
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<p>Thickness of films solidified in slag bulk with different temperatures, slag with basicity 1.28 (<b>a</b>) and 0.85 (<b>b</b>) error bars give standard deviations.</p>
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<p>Thickness of glassy layer of films solidified in slag bulk with different temperatures and solidification times, slag with basicity 1.28 (<b>a</b>) and 0.85 (<b>b</b>), and typical XRD pattern of a solidified slag film (<b>c</b>); the film for XRD measurement was obtained in the high-basicity liquid slag bath with 60 s of probe immersion time.</p>
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<p>Typical total glassy cross section of a low-basicity film solidified in 1400 °C slag and 30 s solidification time, optical image.</p>
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<p>Partial cross section of a low-basicity glassy slag film, bulk slag temperature: 1350 °C, film solidification time: 30 s. Optical image.</p>
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<p>Roughness of film surfaces contacted with copper probe, solidified in slag bulk with different temperatures and solidification times; slag with basicity 1.28 (<b>a</b>) and 0.85 (<b>b</b>).</p>
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<p>Typical surface and profile curve of a high−basicity film (<b>a</b>) and low−basicity film (<b>b</b>), solidified in 1400 °C slag bulk and 30 s solidification time. Backscattered electron images.</p>
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<p>Closed porosity of films solidified in slag bulk with different temperatures and solidification times, slag with basicity 1.28 (<b>a</b>) and 0.85 (<b>b</b>).</p>
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<p>Evolution of thermal conductivities of films upon solidification in molten slag bulk with different temperatures, slag with basicity 1.28 (<b>a</b>) and 0.85 (<b>b</b>).</p>
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16 pages, 6744 KiB  
Article
Annealing Heat Treatment for Homogenizing the Microstructure and Mechanical Properties of Electron-Beam-Welded Thick Plate of Ti-6Al-4V Alloy
by Seongji Seo and Jiyong Park
Materials 2023, 16(23), 7423; https://doi.org/10.3390/ma16237423 - 29 Nov 2023
Cited by 5 | Viewed by 1659
Abstract
In the application of Ti-6Al-4V to aerospace structural components, when welding thick plates similar of the thickness of the components, microstructure and hardness gradients emerge between the base material (BM) and the joint. This leads to the issue of significant stress concentration in [...] Read more.
In the application of Ti-6Al-4V to aerospace structural components, when welding thick plates similar of the thickness of the components, microstructure and hardness gradients emerge between the base material (BM) and the joint. This leads to the issue of significant stress concentration in the BM under tensile stress. To address this problem through post-welding heat treatment, this study conducted heat treatments at temperatures both below (mill annealing, MA) and above the beta-transus temperature (beta annealing, BA) on electron-beam weldments of 18 mm thickness Ti-6Al-4V plates. Subsequently, microstructures and hardness were analyzed at different depths from the upper surface and areas (fusion zone (FZ), heat-affected zone (HAZ), and BM), and tensile properties were measured at various depths. The results indicated that α′ observed in FZ and HAZ was resolved through both MA and BA. Particularly after BA, the microstructural gradient that persisted even after MA completely disappeared, resulting in the homogenization of widmanstätten α + β. Consequently, after BA, the hardness gradient in each zone also disappeared, and the tensile strength was higher than in just-welded and MA heat-treated plates. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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Graphical abstract
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<p>Initial microstructure of Ti-6Al-4V (Ti64) alloy used in this study.</p>
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<p>(<b>a</b>) Schematic diagram of EBW process used in this study and (<b>b</b>) marked area indicating the extracted sample for heat treatment and analysis.</p>
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<p>Tensile sample for used in this study.</p>
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<p>Cross-sectional macrostructural images of (<b>a</b>) EBWed, (<b>b</b>) MA, and (<b>c</b>) BA.</p>
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<p>XRD patterns of A and C zones of the center depth from the upper surface in EBWed and PWHTed Ti64: (<b>a</b>) EBWed Ti64, (<b>b</b>) MA, and (<b>c</b>) BA.</p>
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<p>OM and SEM images for each zone and depth of EBWed with colored arrows that are pink for equiaxed α (α<sub>E</sub>), yellow for grain boundary α (α<sub>GB</sub>), red for α′, orange for widmanstätten α plate (α<sub>w</sub>), blue for β phase, and white for retained β (β<sub>re</sub>); (<b>a</b>–<b>c</b>) OM of zone A (BM), (<b>d</b>–<b>f</b>) OM and (<b>g</b>–<b>i</b>) SEM of zone B (HAZ), (<b>j</b>–<b>l</b>) OM and (<b>m</b>–<b>o</b>) SEM of zone C (FZ).</p>
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<p>OM and SEM images for each zone and depth of MA with colored arrows that are pink for α<sub>E</sub>, yellow for α<sub>GB</sub>, orange for α<sub>w</sub>, and blue for β; (<b>a</b>–<b>c</b>) OM of zone A, (<b>d</b>–<b>f</b>) OM and (<b>g</b>–<b>i</b>) SEM of zone B, (<b>j</b>–<b>l</b>) OM and (<b>m</b>–<b>o</b>) SEM of zone C.</p>
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<p>OM and SEM images at the center for the depths of BA with colored arrows that are yellow for α<sub>GB</sub>, orange for α<sub>w</sub>, and blue for β; (<b>a</b>–<b>f</b>) OM and SEM of zone A, (<b>g</b>–<b>l</b>) OM and SEM of zone B, and (<b>m</b>–<b>r</b>) OM and SEM of zone C.</p>
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<p>Quantitative measurement results of microstructural characteristics for EBWed, MA, and BA according to zones and depths; (<b>a</b>) average grain size, (<b>b</b>) average thickness of α and/or α′ plates.</p>
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<p>Schematic diagram of metallurgical mechanism during EBW and PWHT of Ti64 in this study; (<b>a</b>) before EBWed Ti64, (<b>b</b>) EBWed, (<b>c</b>) MA, and (<b>d</b>) BA.</p>
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<p>Micro-Vickers hardness profiles and measurement areas of EBWed, MA, and BA according to each zone and thickness; (<b>a</b>) hardness measurement areas, (<b>b</b>) 1 mm depth, (<b>c</b>) 9 mm depth, and (<b>d</b>) 17 mm depth.</p>
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<p>Average hardness values of each zone in EBWed, MA, and BA according to depth.</p>
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<p>Tensile properties of EBWed, MA, and BA according to depth; (<b>a</b>) tensile strength and (<b>b</b>) elongation.</p>
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<p>Tensile fractured samples and fractographies of EBWed, MA, and BA according to depths; (<b>a</b>–<b>c</b>) EBWed, (<b>d</b>–<b>f</b>) MA, and (<b>g</b>–<b>i</b>) BA.</p>
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19 pages, 6181 KiB  
Article
The Effect of Mn and Ti Ratio on Microstructure and Mechanical and Machinability Properties of 316 L Stainless Steel Used in Biomedical Applications
by Mustafa Türkmen, Alla Muhammed Tanouz, Mahir Akgün and Mehmet Akif Erden
Metals 2023, 13(11), 1804; https://doi.org/10.3390/met13111804 - 26 Oct 2023
Cited by 2 | Viewed by 1576
Abstract
In this study, titanium (Ti) and manganese (Mn) element powders in determined amounts (0.35–0.75 and 1.5 wt %) were added into the 316 L stainless steel matrix by means of powder metallurgy (PM) technology, either individually or in pairs, and the desired composition [...] Read more.
In this study, titanium (Ti) and manganese (Mn) element powders in determined amounts (0.35–0.75 and 1.5 wt %) were added into the 316 L stainless steel matrix by means of powder metallurgy (PM) technology, either individually or in pairs, and the desired composition was obtained as a powder mixture. The powders used in the study were cold-pressed tensile sample molds prepared in ASTM E8M standards, unidirectionally cold-pressed under 750 MPa compression pressure and formed into blocks. After pressing, the raw strength samples were sintered in an atmosphere-controlled tube furnace at 1250 °C for two hours in an argon atmosphere. The microstructure and mechanical properties of the produced PM steels were characterized using an optical microscope, SEM, EDS, tensile test, and hardness test. The results showed that the stainless steel samples with 0.35 (Ti and Mn) added to 316 L stainless steel had the highest yield strength, tensile strengths, and hardness strengths. However, with the addition of 0.75–1.5 Ti, 0.75–1.5 Mn and 0.75–1.5 (Ti and Mn) to 316 L stainless steel, a decrease was observed in the mechanical properties. Moreover, the stainless steel sample with 0.35 (Ti and Mn) added to 316 L stainless steel is better than other samples in terms of surface quality. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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<p>Schematic representation of powder metallurgy production stages.</p>
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<p>Schematic representation of milling experiments.</p>
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<p>Optical microscope images of samples produced by the PM method before and after adding Ti and Mn (500×).</p>
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<p>SEM and spot EDS results from Alloy I.</p>
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<p>SEM image and EDS points for Alloy X.</p>
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<p>SEM image and EDS for Alloy I (<b>a</b>) and Alloy X (<b>b</b>).</p>
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<p>Tensile curves of 316 L PM steels containing different ratios of Ti-Nb (a—16 L; b—16 L + 0.35 Ti; c—16 L + 0.35 Mn; d—16 L + 0.35 (Ti + Mn); e—16 L + 0.75 Ti; f—16 L + 0.75 Mn; g—16 L + 0.75 (Ti + Mn); h—16 L + 1.5 Ti; i—16 L + 1.5 Mn; j—16 L + 1.5 (Ti + Mn)).</p>
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<p>Fractured surface images of samples (100×–400×).</p>
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<p>Fractured surface images of samples (100×–400×).</p>
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<p>Fractured surface images of samples (100×–400×).</p>
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<p>Fracture surface (1000×) and point EDS results for Alloy VIII.</p>
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<p>Surface roughness assessment.</p>
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<p>Cutting temperature assessment.</p>
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14 pages, 8160 KiB  
Article
Formation of Cu Nanotwins on Silicon Carbide Wafers with Cr Adhesive Layer under Various Substrate Bias
by Devi Indrawati Syafei, Meng-Ting Chiang and Tung-Han Chuang
Metals 2023, 13(10), 1747; https://doi.org/10.3390/met13101747 - 15 Oct 2023
Cited by 2 | Viewed by 1459
Abstract
This study focuses on the analyses of nano-twinned copper (Cu) films deposited through magnetron sputtering on silicon carbide (SiC) chips. The investigation encompasses the utilization of a chromium (Cr) adhesive layer coupled with varying voltage bias conditions. The goal is to comprehensively examine [...] Read more.
This study focuses on the analyses of nano-twinned copper (Cu) films deposited through magnetron sputtering on silicon carbide (SiC) chips. The investigation encompasses the utilization of a chromium (Cr) adhesive layer coupled with varying voltage bias conditions. The goal is to comprehensively examine the influence of the adhesive layer and negative bias voltages, contributing to an enhanced understanding of materials engineering and bonding technologies for advanced applications. The formation of a nano-twinned structure and (111) surface orientation can be properly controlled by applied substrate bias. High-density nanotwinned structures were introduced into Cu films sputtered on SiC substrates with 82.3% of (111) orientation proportion at −150 V, much higher than the Cu film sputtered with another substrate bias. It is concluded that the sputtered Cu nanotwinned film formed with −150 V bias voltage has the potential to be employed as the interlayer for low-temperature direct bonding. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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<p>Top view of FIB metallography of sputtered SiC/Cr/nt-Cu with different biases; (<b>a</b>) −100 V, (<b>b</b>) −150 V, (<b>c</b>) −200 V, and (<b>d</b>) −250 V.</p>
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<p>FIB metallography of the cross-section of sputtered SiC/Cr/nt-Cu with different biases; (<b>a</b>) −100 V, (<b>b</b>) −150 V, (<b>c</b>) −200 V, and (<b>d</b>) −250 V.</p>
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<p>XRD spectra of Cu nanotwinned films on Cr pre-coated SiC substrates with various biases.</p>
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<p>Top-view EBSD inverse pole figure (IPF) maps of sputtered SiC/Cr/nt-Cu films deposited with substrate biases of (<b>a</b>) −100 V, (<b>b</b>) −150 V, (<b>c</b>) −200 V, and (<b>d</b>) −250 V.</p>
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<p>Plane-view EBSD inverse pole figure (IPF) maps of sputtered Cu films deposited with substrate biases of (<b>a</b>) −100 V, (<b>b</b>) −150 V, (<b>c</b>) −200 V, and (<b>d</b>) −250 V showing (111) orientation.</p>
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<p>EBSD analyses of the cross-section of sputtered Cu nanotwinned films on SiC substrate with the Cr pre-coating EBSD inverse pole figure mapping (<b>a</b>) −150 V; (<b>b</b>) −200 V; EBSD (111) grain mapping (<b>c</b>) −150 V; (<b>d</b>) −200 V.</p>
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<p>Cross-sectional TEM micrographs of as-deposited Cu nanotwinned film deposited with −150 V bias: (<b>a</b>) bright-field TEM image of columnar Cu nanotwin, (<b>b</b>) diffraction pattern taken along the [110] zone axis, and (<b>c</b>) high-magnification HRTEM image of nanotwinned structures with different fast Fourier transformation (FFT) patterns of twin and matrix.</p>
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<p>(<b>a</b>) Cross-sectional bright-field TEM image of area between columnar grains and (<b>b</b>) distribution of twin thicknesses in the highly (111)-oriented Cu film.</p>
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<p>(<b>a</b>,<b>b</b>) Cross-sectional bright-field TEM image reveals the presence of fine lamellar cells at the interface between the nanotwinned columnar grains.</p>
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15 pages, 4238 KiB  
Article
Effect of Furnace Structure on Burden Distribution and Gas Flow in Sinter Vertical Cooling Furnace
by Haifeng Li, Tengfei Qi and Yongjie Zhang
Appl. Sci. 2023, 13(20), 11268; https://doi.org/10.3390/app132011268 - 13 Oct 2023
Cited by 7 | Viewed by 1306
Abstract
Sinter sensible heat recovery via a vertical cooling furnace is a new type of waste heat recovery process proposed based on coke dry quenching. However, the segregation of the burden in a vertical cooling furnace is serious, resulting in a large amount of [...] Read more.
Sinter sensible heat recovery via a vertical cooling furnace is a new type of waste heat recovery process proposed based on coke dry quenching. However, the segregation of the burden in a vertical cooling furnace is serious, resulting in a large amount of cooling gas escaping from the short-circuit channel of the vertical cooling furnace, which seriously affects the uniform gas–solid heat transfer in the furnace. To improve the burden distribution and gas flow in such a furnace, this paper proposes a Venturi-type vertical cooling furnace. Based on the single silo of a vertical cooling furnace in Meishan Steel, a slot model was established, and the improvement effect of the Venturi furnace structure on the burden distribution and gas flow was studied using the DEM–CFD coupling method. The results show that compared with the existing furnace type, the inclined wall of the Venturi furnace changed the direction of the high Dnv (average diameter) channel from vertical to inclined-vertical and reduced the Dnv from >0.033 m to 0.028~0.03 m in the vertical part of the variable-diameter section, thus reducing the influence area of the high Dnv channel. The minimum and average values of the voidage in the contraction part of the variable-diameter section increased from 0.28 and 0.315 to 0.31 and 0.33, respectively, which caused the voidage distribution to change from U-shaped to W-shaped along the longitudinal direction while simultaneously reducing the longitudinal fluctuation range of the voidage from 0.28~0.39 to 0.298~0.37. The gas flow direction changed from vertical-upward to vertical-inclined-upward, which increased the gas–solid contact. The gas velocity increased significantly. In the vertical section, the average gas velocity was 2.34 m/s, which was 30.73% higher than the velocity of 1.79 m/s of the existing furnace type. In the variable-diameter section, the average gas velocity was 3.52 m/s, which was 72.55% higher than the velocity of 2.04 m/s of the existing furnace type. The high-speed gas channel basically only existed in the sidewall area and the center area of the vertical section, and the length was reduced from 3.11 m to 2.52 m, which reduced the influence area. In the variable-diameter section, the high-speed gas channel disappeared, and the uniformity of the gas velocity distribution was greatly improved. The gas pressure drop increased from 4140 Pa to 6410 Pa, with an increase of 54.83%. Therefore, when designing the Venturi furnace type, it was necessary to take into consideration the improvement in the gas velocity distribution and the increase in the pressure drop. The research results of this paper can provide guidance for the structure optimization of the sinter vertical cooling furnace. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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<p>Existing furnace type and Venturi furnace type. (<b>a</b>) Model dimension. (<b>b</b>) Slot model. (<b>c</b>) Computational grid.</p>
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<p>Influence of gas velocity on pressure drop.</p>
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<p>Particle distribution during the discharging process. (<b>a</b>) Existing furnace type. (<b>b</b>) Venturi furnace type.</p>
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<p>Distribution of particle size and <span class="html-italic">D</span><sub>nv</sub> in the existing furnace type and Venturi furnace type. (<b>a</b>) Particle size distribution. (<b>b</b>) <span class="html-italic">D</span><sub>nv</sub> distribution.</p>
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<p>Voidage distribution of sinter layer. (<b>a</b>) Voidage distribution map. (<b>b</b>) Voidage distribution at line V1.</p>
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<p>Gas velocity vector distribution. (<b>a</b>) Existing furnace type. (<b>b</b>) Venturi furnace type.</p>
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<p>Gas velocity distribution. (<b>a</b>) Existing furnace type. (<b>b</b>) Venturi furnace type.</p>
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<p>High-speed gas channel distribution in different furnaces. (<b>a</b>) Existing furnace type. (<b>b</b>) Venturi furnace type and. (<b>c</b>) Venturi furnace type.</p>
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<p>Gas pressure distribution. (<b>a</b>) Existing furnace type. (<b>b</b>) Venturi furnace type.</p>
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28 pages, 6710 KiB  
Article
The Importance of Mixing Time in Intensely Stirred Metallurgical Reactors: Applied on Decarburization Reactions
by Serg Chanouian, Henrik Larsson and Mikael Ersson
Metals 2023, 13(10), 1694; https://doi.org/10.3390/met13101694 - 5 Oct 2023
Viewed by 1283
Abstract
In metallurgical converter processes, numerical modeling is a useful tool for understanding the complexity of the systems. In this paper, we present a practical model that couples fluid dynamics and chemical reactions to explore the impact of mixing time on decarburization. Using computational [...] Read more.
In metallurgical converter processes, numerical modeling is a useful tool for understanding the complexity of the systems. In this paper, we present a practical model that couples fluid dynamics and chemical reactions to explore the impact of mixing time on decarburization. Using computational fluid dynamics (CFD), in this study, we investigate an arbitrary metallurgical reactor with continuous oxygen supply, focusing on the Fe–C–O system. The model employs local equilibrium, a turbulence limiter, and finite volume method for mass, momentum, and energy transfer. Tracer injection points in the gas plume’s rising region exhibit faster mixing, and a comparison of reaction cases reveals distinct decarburization rates based on oxygen injection distribution and the influence of turbulence on reactions. Overall, while mixing time matters, the results show that this system is primarily governed by thermodynamics and oxygen supply, and a 270% increase in mixing time increase had a small impact on the end carbon content. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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<p>Schematic illustration of numerical geometry, side and top views.</p>
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<p>Image capturing the penetration depth of the physical water model. The red-dashed box represents the injection zone utilized in the numerical model.</p>
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<p>Flowchart of the fluid transport and chemical reactions coupling scheme.</p>
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<p>(<b>a</b>) Illustration of the difference between the control volume and control mass approach; (<b>b</b>) illustration of the mass conservation with constant density where total mass is conserved.</p>
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<p>Figure illustrating 9 different positions of tracer injections with the red dots.</p>
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<p>Schematic illustration of turbulent reaction rate treatment in coupled model, where m denotes the mass and T is the temperature in the cell.</p>
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<p>Mixing time comparison between the numerical model and the water model. (<b>a</b>) Shows the average value of mixing time based on all tracer injections which are displayed in (<b>b</b>).</p>
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<p>Displays the velocity magnitude and circular flow field caused by the rising gas in iso-planes. Velocity vectors and streamlines illustrate the flow direction.</p>
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<p>Mixing times for different tracer injections with varying diffusion coefficients.</p>
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<p>The graphs show the element and temperature evolution in the domain over time: (<b>a</b>) Carbon mass; (<b>b</b>) iron mass; (<b>c</b>) oxygen mass; (<b>d</b>) temperature.</p>
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<p>The graphs show the element and temperature evolution in the domain over time: (<b>a</b>) Carbon mass; (<b>b</b>) iron mass; (<b>c</b>) oxygen mass; (<b>d</b>) temperature.</p>
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<p>Iso-planes display the carbon content at the end of the simulation (660 s) in Cases 1a–4a. Note, there is no carbon oxidation in the above region of Case 2a, the gradient in the upper iso-planes is the effect of the flow spreading the reacted melt.</p>
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<p>Iso-planes display the equilibrium fraction due to turbulence mixing in the cells for Case 4. The transparent black volume (plume shape) illustrates the reaction zone. Note that values higher than one are displayed as 1.</p>
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<p>The element mass in the domain over time using effective diffusion (Cases 1a, 2a, and 4a) as well as only laminar diffusion (Cases 1b, 2b, and 4b): (<b>a</b>) Carbon mass; (<b>b</b>) iron mass.</p>
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<p>Four time frames between 0 and 7 s, showing the local changes of carbon content in Case 4a (effective diffusion) and 4b (laminar diffusion only).</p>
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<p>Compares the maximum, average, and minimum carbon content in the melt over 20 s for Cases 4a (including turbulent diffusion) and 4b (laminar diffusion only).</p>
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<p>The graphs show the mass of carbon in the domain over time for Cases 1–4 between Test a (higher initial oxygen content) and Test c: (<b>a</b>) Case 1; (<b>b</b>) Case 2; (<b>c</b>) Case 3; (<b>d</b>) Case 4.</p>
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<p>The graphs show the mass of carbon in the domain over time for Cases 1–4 between Test a (higher initial oxygen content) and Test c: (<b>a</b>) Case 1; (<b>b</b>) Case 2; (<b>c</b>) Case 3; (<b>d</b>) Case 4.</p>
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<p>The graphs show the mass of oxygen in the domain over time for Case 4 between Test a (higher initial oxygen content) and Test c.</p>
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30 pages, 3282 KiB  
Article
Steel, Aluminum, and FRP-Composites: The Race to Zero Carbon Emissions
by Vaishnavi Vijay Rajulwar, Tetiana Shyrokykh, Robert Stirling, Tova Jarnerud, Yuri Korobeinikov, Sudip Bose, Basudev Bhattacharya, Debashish Bhattacharjee and Seetharaman Sridhar
Energies 2023, 16(19), 6904; https://doi.org/10.3390/en16196904 - 30 Sep 2023
Cited by 15 | Viewed by 4672
Abstract
As various regions around the world implement carbon taxes, we assert that the competitiveness of steel products in the marketplace will shift according to individual manufacturers’ ability to reduce CO2 emissions as measured by cradle-to-gate Life Cycle Analysis (LCA). This study was [...] Read more.
As various regions around the world implement carbon taxes, we assert that the competitiveness of steel products in the marketplace will shift according to individual manufacturers’ ability to reduce CO2 emissions as measured by cradle-to-gate Life Cycle Analysis (LCA). This study was performed by using LCA and cost estimate research to compare the CO2 emissions and the additional cost applied to the production of various decarbonized materials used in sheet for automotive industry applications using the bending stiffness-based weight reduction factor. The pre-pandemic year 2019 was used as a baseline for cost estimates. This paper discusses the future cost scenarios based on carbon taxes and hydrogen cost. The pathways to decarbonize steel and alternative materials such as aluminum and reinforced polymer composites were evaluated. Normalized global warming potential (nGWP) estimates were calculated assuming inputs from the current USA electricity grid, and a hypothetical renewables-based grid. For a current electricity grid mix in the US (with 61% fossil fuels, 19% nuclear, 20% renewables), the lowest nGWP was found to be secondary aluminum and 100% recycled scrap melting of steel. This is followed by the natural gas Direct Reduced Iron–Electric Arc Furnace (DRI-EAF) route with carbon capture and the Blast Furnace-Basic Oxygen Furnace (BF-BOF) route with carbon capture. From the cost point of view, the current cheapest decarbonized production route is natural gas DRI-EAF with Carbon Capture and Storage (CCS). For a renewable electricity grid (50% solar photovoltaic and 50% wind), the lowest GWP was found to be 100% recycled scrap melting of steel and secondary aluminum. This is followed by the hydrogen-based DRI-EAF route and natural gas DRI-EAF with carbon capture. The results indicate that, when applying technologies available today, decarbonized steel will remain competitive, at least in the context of automotive sheet selection compared to aluminum and composites. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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<p>BF-BOF pathway and CO<sub>2</sub> emissions (based on—IEA, 2013 [<a href="#B36-energies-16-06904" class="html-bibr">36</a>]) (green stands for renewable grid).</p>
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<p>NG-DRI and EAF steelmaking (green stands for renewable grid).</p>
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<p>Primary Al production route (green stands for renewable grid).</p>
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<p>A schematic overview of glass-fiber-reinforced composites production (green stands for renewable grid)<b>.</b> * The multistage resin production process is described by Hill et al. [<a href="#B54-energies-16-06904" class="html-bibr">54</a>]. ** Considering fabrication by SMC.</p>
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<p>Schematic overview of carbon-fiber-reinforced composite production. * Multistage resin production process is described by Hill et al. [<a href="#B54-energies-16-06904" class="html-bibr">54</a>]. ** Considering fabrication by SMC.</p>
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<p>Manufacturing cost, carbon tax applied, and Global Warming Potential for the studied materials and processes (traditional grid). These are displayed per tonne of finished sheet of the material at factory gate. The carbon tax applied is USD 80 per tonne of CO<sub>2</sub> equivalent emissions as measured using the cradle-to-gate LCA method. The LCA assumes a traditional electricity grid with a GWP of 0.385 kg CO<sub>2 eq</sub>/kWh. Note the break in the manufacturing cost axis as carbon fiber composite is more expensive per tonne than other materials.</p>
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<p>Normalized manufacturing cost, carbon tax, and Global Warming Potential for the studied materials and processes (traditional grid). The masses of the materials are normalized using elastic modulus and density as described in Methods to result in functionally equivalent sheet products. The LCA assumes a traditional electricity grid with GWP of 0.385 kg CO<sub>2 eq</sub>/kWh.</p>
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<p>Manufacturing cost, carbon tax applied, and Global Warming Potential for the studied materials and processes (renewables-based grid). These are displayed per tonne of finished sheet of the material at factory gate. The carbon tax applied is USD 80 per tonne of CO<sub>2</sub> equivalent emissions as measured using the cradle-to-gate LCA method. The LCA assumes a renewables-based electricity grid with a GWP of 0.038 kg CO<sub>2 eq</sub>/kWh. Note the break in the manufacturing cost axis as carbon fiber composite is more expensive per tonne than other materials.</p>
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<p>Normalized manufacturing cost, carbon tax, and Global Warming Potential for the studied materials and processes (renewables-based grid). The masses of the materials are normalized using elastic modulus and density as described in Methods to result in functionally equivalent sheet products. The LCA assumes a renewables-based electricity grid with GWP of 0.038 kg CO<sub>2 eq</sub>/kWh.</p>
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<p>Sensitivity analysis of carbon tax on total manufacturing costs (USD/t steel sheet <sub>eq</sub>) (traditional grid). Total cost per tonne is defined as the total manufacturing cost of the material plus applied carbon taxes. The costs are normalized according to the equivalent mass method for a sheet product as described in Methods. The midline of each bar depicts the base case of a USD 80/t CO<sub>2</sub> tax applied on carbon emissions. The left boundary of the gray bar depicts the manufacturing cost with no carbon tax applied. The right boundary of the yellow bar depicts the application of a carbon tax equal to USD 160/t CO<sub>2</sub>. The LCA assumes a traditional electricity grid with a GWP of 0.385 kg CO<sub>2 eq</sub>/kWh.</p>
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<p>Sensitivity analysis of carbon tax on total manufacturing costs (USD/t steel sheet <sub>eq</sub>) (renewables-based grid). Total cost per tonne is defined as the total manufacturing cost of the material plus applied carbon taxes. The costs are normalized according to the equivalent mass method for a sheet product as described in Methods. The midline of each bar depicts the base case of USD 80/t CO<sub>2</sub> tax applied on carbon emissions. The left boundary of the gray bar depicts the manufacturing cost with no carbon tax applied. The right boundary of the yellow bar depicts the application of a carbon tax equal to USD 160/t CO<sub>2</sub>. The LCA assumes a renewables-based electricity grid with a GWP of 0.038 kg CO<sub>2 eq</sub>/kWh.</p>
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<p>Sensitivity analysis of hydrogen-based and natural gas-based DRI technologies to feedstock cost.</p>
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12 pages, 4535 KiB  
Article
An Experiment on Surface Fluctuation of Ga-In-Sn Alloy with a Permanent Magnet Flow Control Mold
by Zefeng Han, Engang Wang, Zhongxin Zhai and Zepeng Wang
Metals 2023, 13(10), 1662; https://doi.org/10.3390/met13101662 - 28 Sep 2023
Viewed by 997
Abstract
To control well the surface fluctuation of liquid metal in a slab mold, a new type of combined permanent magnets braking system, namely a permanent magnet flow control mold (PMFC-Mold) is proposed by our research group, of which its main feature is that [...] Read more.
To control well the surface fluctuation of liquid metal in a slab mold, a new type of combined permanent magnets braking system, namely a permanent magnet flow control mold (PMFC-Mold) is proposed by our research group, of which its main feature is that the device can control the flow of molten steel in the mold without additional energy. To observe the fluctuation state of the alloy with the PMFC-Mold, instantaneous surface fluctuations were recorded by a laser level meter and camera. To study the effect of various casting speeds and permanent magnet placement on surface fluctuations, the three measurement points, which were 7, 18, and 36 mm away from the narrow surface of the mold, were selected to record the trend of level fluctuation. Three types of permanent magnet placement were designed by setting the differences between the height center of the permanent magnet and the free surface in the slab mold, which were H1 = 0 mm, H2 = −25 mm, and H3 = −75 mm. The experimental results indicated that with the acceleration of the casting speed, the average height and standard deviation of surface fluctuation at the measurement point increased, but the surface fluctuation pattern remained. When the permanent magnets were arranged at H1 = 0 mm and H2 = −25 mm, the position of the magnetic field was reasonable and the surface fluctuation could be effectively suppressed. In contrast, when the permanent magnets were arranged at H3 = −75 mm, the level fluctuation was intensified. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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<p>Schematic diagram of Ga-In-Sn alloy magnetic fluid experimental platform and platform structure.</p>
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<p>Schematic diagram of permanent magnets combination device, permanent magnets arrangement (arrow pointing to N pole) and measurement points location.</p>
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<p>The relative position between the height center of PMFC-Mold permanent magnet and the free liquid surface, (<b>a</b>) H<sub>1</sub> = 0 mm, (<b>b</b>) H<sub>2</sub> = −25 mm, (<b>c</b>) H<sub>3</sub> = −75 mm.</p>
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<p>The characteristic lines and the magnetic field induction intensity distribution on the corresponding characteristic lines: (<b>a</b>) measurement position of characteristic lines; (<b>b</b>) magnetic flux density along the X direction; (<b>c</b>) magnetic flux density along the Y direction; (<b>d</b>) magnetic flux density along the Z direction.</p>
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<p>Experimental data of instantaneous level fluctuation under different casting speeds at 3 measuring points: (<b>a</b>) Measuring point 1 (distance from narrow surface L = 7 mm); (<b>b</b>) Measuring point 2 (distance from narrow surface L = 18 mm); and (<b>c</b>) Measuring point 3 (distance from narrow surface L = 36 mm).</p>
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<p>Arithmetic mean <span class="html-italic">ā</span> (<b>a</b>) and standard deviation <span class="html-italic">α</span> (<b>b</b>) of level fluctuation at different measurement points.</p>
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<p>Real time data collection of level fluctuations for PMFC-Mold as permanent magnets arranged in different positions.</p>
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<p>Arithmetic mean <span class="html-italic">ā</span> (<b>a</b>) and standard deviation <span class="html-italic">α</span> (<b>b</b>) of level fluctuation for PMFC-Mold permanent magnets arranged at different positions.</p>
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<p>Diagram of level fluctuation in the wide surface direction of the mold.</p>
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<p>Diagram of level fluctuation of free surface in the mold.</p>
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17 pages, 4109 KiB  
Article
Research and Application of Coupled Mechanism and Data-Driven Prediction of Blast Furnace Permeability Index
by Kangkang Tan, Zezheng Li, Yang Han, Xiwei Qi and Wei Wang
Appl. Sci. 2023, 13(17), 9556; https://doi.org/10.3390/app13179556 - 23 Aug 2023
Cited by 4 | Viewed by 1910
Abstract
In order to ensure the stable operation of blast furnace production, it is necessary to keep abreast of the trends in the gas permeability index of the blast furnace. As one of the key parameters to be monitored in the process of blast [...] Read more.
In order to ensure the stable operation of blast furnace production, it is necessary to keep abreast of the trends in the gas permeability index of the blast furnace. As one of the key parameters to be monitored in the process of blast furnace smelting, the gas permeability index directly reflects the performance of the blast furnace in the actual production of the furnace. Continuous monitoring of the permeability index is required in the actual production of the blast furnace in order to effectively guarantee the stable and smooth operation of the blast furnace. The aim of this study is to accurately predict the trend in the blast furnace gas permeability index by constructing an intelligent prediction model and utilizing a data-driven approach to monitor the gas permeability index and ensure the stable operation of the blast furnace. First, based on the actual production data of a #2 blast furnace of an iron and steel enterprise, an isolated forest algorithm is applied to detect and eliminate the outliers in the original data, and then a data driver set is constructed after normalization of the deviation. Second, by analyzing the coupling mechanism between the blast furnace permeability and gas flow, as well as Spearman correlation analysis and MIC maximum information coefficient (MIC) analysis, key parameters are screened out as feature variables from the data-driven set. Finally, a wavelet neural network algorithm is used to construct an intelligent prediction model of the blast furnace gas permeability index. Compared with a BP neural network (BP), a particle swarm-optimized BP neural network (PSO-BP), and XGBoost, the wavelet neural network shows obvious advantages when the error is controlled in the range of ±0.1, and the prediction accuracy can reach 95.71%. The model is applied to the actual production of a #2 blast furnace of an iron and steel enterprise, and the results show that the predicted value of the blast furnace permeability index is highly consistent with the actual value of real-time blast furnace production, which verifies its excellent characteristics. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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<p>Isolated forest structure graph.</p>
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<p>Blast furnace structure.</p>
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<p>Blast furnace permeability-related factors.</p>
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<p>Blast furnace gas flow influence factors.</p>
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<p>Heat map of Spearman’s characteristic correlation analysis.</p>
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<p>MIC maximum information coefficient.</p>
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<p>Wavelet neural network structure.</p>
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<p>Morlet wavelet function diagram.</p>
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<p>Flow chart of wavelet neural network prediction steps.</p>
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<p>Wavelet neural network model blast furnace permeability index prediction graph.</p>
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<p>Comparison of predicted values and actual production values.</p>
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15 pages, 5209 KiB  
Article
The Influence of Injection Temperature and Pressure on Pattern Wax Fluidity
by Viacheslav E. Bazhenov, Andrey V. Sannikov, Elena P. Kovyshkina, Andrey V. Koltygin, Andrey I. Bazlov, Vladimir D. Belov and Dmitry N. Dmitriev
J. Manuf. Mater. Process. 2023, 7(4), 141; https://doi.org/10.3390/jmmp7040141 - 4 Aug 2023
Cited by 3 | Viewed by 1985
Abstract
In the investment casting process, the pattern made of wax is obtained in a die for further formation of a shell mold. The problem of die-filling by pattern wax is significant because it influences the quality of the final casting. This work investigates [...] Read more.
In the investment casting process, the pattern made of wax is obtained in a die for further formation of a shell mold. The problem of die-filling by pattern wax is significant because it influences the quality of the final casting. This work investigates three commercial pattern waxes’ fluidity with a newly developed injection fluidity test. It was shown that the fluidity of waxes increased with increasing injection temperature and pressure, and the simultaneous increase in temperature and pressure gives a much more significant enhancement of fluidity than an increase in temperature or pressure separately. The rheological behavior of the waxes was also investigated at different temperatures using a rotational viscosimeter, and temperature dependences of waxes’ dynamic viscosity were determined. It was shown that wax viscosity is increased more than ten times with decreasing temperature from 90 to 60 °C. A good correlation between wax fluidity and its viscosity is observed, which is different from metallic alloys, where the solidification behavior is more critical. The difference in wax flow behavior in comparison with metallic melts is associated with the difference in dynamic viscosity, which for investigated waxes and metallic melts is 3000–27,000 mPa·s and 0.5–6.5 mPa·s, respectively. The difference in investigated filled waxes’ fluidity is observed, which can be associated with the type and amount of filler. The twice-increasing fraction of cross-linked polystyrene decreases fluidity twice. At the same time, terephthalic acid has a minor influence on wax fluidity. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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<p>The die used for fluidity test: (<b>a</b>) top view and cross-section and (<b>b</b>) probe and die in 3D.</p>
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<p>Typical fluidity probes of (<b>a</b>,<b>d</b>,<b>g</b>,<b>j</b>) RG20, (<b>b</b>,<b>e</b>,<b>h</b>,<b>k</b>) S1235, and (<b>c</b>,<b>f</b>,<b>i</b>,<b>l</b>) S1135 waxes at (<b>a</b>–<b>f</b>) 70 and (<b>g</b>–<b>l</b>) 90 °C injection temperatures and (<b>a</b>–<b>c</b>,<b>g</b>–<b>i</b>) 49 and (<b>d</b>–<b>f</b>,<b>j</b>–<b>l</b>) 196 kPa injection pressures.</p>
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<p>Fluidity of (<b>a</b>) RG20, (<b>b</b>) S1235, and (<b>c</b>) S1135 waxes at various injection temperatures and pressures.</p>
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<p>Contour plot showing the influence of injection temperatures and pressures on the fluidity of (<b>a</b>) RG20, (<b>b</b>) S1235, and (<b>c</b>) S1135 waxes.</p>
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<p>Fluidity of RG20, S1235, and S1135 waxes depending on injection temperatures at pressures (<b>a</b>) <span class="html-italic">P</span> = 49 kPa, (<b>b</b>) <span class="html-italic">P</span> = 98 kPa, (<b>c</b>) <span class="html-italic">P</span> = 147 kPa, and (<b>d</b>) <span class="html-italic">P</span> = 196 kPa.</p>
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<p>The influence of rotational frequency on the viscosity of (<b>a</b>) RG20, (<b>b</b>) S1235, and (<b>c</b>) S1135 waxes at various temperatures and (<b>d</b>) mean viscosity at various temperatures of waxes.</p>
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<p>The influence of rotational frequency on the viscosity of (<b>a</b>) RG20, (<b>b</b>) S1235, and (<b>c</b>) S1135 waxes at various temperatures and (<b>d</b>) mean viscosity at various temperatures of waxes.</p>
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<p>The heating and cooling DSC curves for RG20, S1235, and S1135 waxes.</p>
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<p>The influence of RG20, S1235, and S1135 waxes viscosity on its fluidity at different injection pressures. Each fluidity point corresponds to the viscosity at the same temperature.</p>
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23 pages, 6456 KiB  
Article
Comparative Analysis of the Hot Isostatic Pressing Densification Behavior of Uniform and Non-Uniform Distributed Powder
by Fandi Meng, Lihui Lang and Yi Xiao
Metals 2023, 13(7), 1319; https://doi.org/10.3390/met13071319 - 24 Jul 2023
Cited by 4 | Viewed by 2425
Abstract
Hot isostatic pressing (HIP) technology can directly produce nearly clean shaped workpieces that meet the requirements while ensuring machining accuracy and surface quality. Usually, people use numerical simulation methods to reduce experimental costs. Generally, a uniform powder relative density distribution of about 65% [...] Read more.
Hot isostatic pressing (HIP) technology can directly produce nearly clean shaped workpieces that meet the requirements while ensuring machining accuracy and surface quality. Usually, people use numerical simulation methods to reduce experimental costs. Generally, a uniform powder relative density distribution of about 65% is used in the simulation. However, in practical engineering, we found that even with additional tools such as vibration tables, the powder filling is not uniform. The non-uniform distribution causes uneven shrinkage of the powder and capsule after HIP. In this paper, a numerical model for HIPing of Ti-6Al-4V powder is developed to improve the prediction by comparing the uniform and non-uniform initial powder distribution. The results show that different initial relative density distributions affect the powder densification process and further affect the deformation of the capsule. It also leads to non-uniform stress distribution after HIP, which increases the risk of capsule rupture. The analysis of the numerical simulation results and the comparison with the experimental results highlights that taking into account the non-uniform powder distribution inside the capsule is vital to improve numerical results and produce near-net shape components. The maximum error of the simulation with the usual initial relative density setting of 65% is 4.2%. However, considering the uneven distribution of initial powder, the maximum error is reduced to 3.16%, and the average error is also less than 2%. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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<p>Structural diagram of the capsule and welded one.</p>
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<p>Powder particle morphology.</p>
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<p>HIP process route.</p>
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<p>Schematic diagram of grid model and numerical analysis points. A-G are the sampling sites.</p>
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<p>Displacement vector distributions of model I at different times. (<b>a</b>) Displacement vector distributions at 30 min; (<b>b</b>) Displacement vector distributions at 2 h; (<b>c</b>) Displacement vector distributions at 5 h; (<b>d</b>) Displacement vector distributions at 7 h.</p>
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<p>Displacement vector distributions of model II at different times. (<b>a</b>) Displacement vector distributions at 30 min; (<b>b</b>) Displacement vector distributions at 2 h; (<b>c</b>) Displacement vector distributions at 5 h; (<b>d</b>) Displacement vector distributions at 7 h.</p>
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<p>Displacement histories of model I and model II at different locations of test line L1. (<b>a</b>) Displacement histories of model I. (<b>b</b>) Displacement histories of model II.</p>
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<p>Radial displacement of simulation I and II. (<b>a</b>) simulation I; (<b>b</b>) simulation II.</p>
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<p>Displacement of histories of simulation II.</p>
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<p>Schematic diagram of dimension measurement after HIP.</p>
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<p>Relative density distributions of model I at different times. (<b>a</b>) Relative density distributions at 30 min; (<b>b</b>) Relative density distributions at 1 h; (<b>c</b>) Relative density distributions at 4 h; (<b>d</b>) Relative density distributions at 7 h.</p>
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<p>Relative density distributions of model II at different times. (<b>a</b>) Relative density distributions at 30 min; (<b>b</b>) Relative density distributions at 1 h; (<b>c</b>) Relative density distributions at 4 h; (<b>d</b>) Relative density distributions at 7 h.</p>
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<p>Density change process at test line L1 in model I and II. (<b>a</b>) Density change process of model I; (<b>b</b>) Density change process of model II.</p>
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<p>Variation trends of the relative densities of test line L2 of model I.</p>
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<p>Variation trends of the relative densities of test line L2 of model II.</p>
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<p>Comparison of the overall densification trends of model I and model II.</p>
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<p>The Mises stress distribution of the powder: (<b>a</b>) Uniform initial density distribution, (<b>b</b>) non-uniform initial density.</p>
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<p>The stress distribution of the capsule: (<b>a</b>) uniform initial density distribution, (<b>b</b>) non-uniform initial density. A–C are the sampling points.</p>
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<p>Variation diagram of equivalent Cauchy stress of each node. (<b>a</b>) Uniform initial density; (<b>b</b>) Non-uniform initial density.</p>
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<p>Schematic diagram of sampling location. A–C are the sampling sites.</p>
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<p>Microstructure of A part.</p>
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15 pages, 4029 KiB  
Article
Cooling Rate Modeling and Evaluation during Centrifugal Atomization Process
by Sasha A. Cegarra, Jordi Pijuan and María D. Riera
J. Manuf. Mater. Process. 2023, 7(3), 112; https://doi.org/10.3390/jmmp7030112 - 7 Jun 2023
Cited by 4 | Viewed by 2832
Abstract
Centrifugal atomization is a rapid solidification technique involving fast cooling rates to produce high-quality powders. The final microstructure of the atomized particles is closely linked with the thermal history and cooling rates experienced during the atomization process. However, there is insufficient research on [...] Read more.
Centrifugal atomization is a rapid solidification technique involving fast cooling rates to produce high-quality powders. The final microstructure of the atomized particles is closely linked with the thermal history and cooling rates experienced during the atomization process. However, there is insufficient research on the temperature evolution of metal particles produced by this technique, and most works evaluate the thermal history of the droplet through semi-empirical correlations that lie outside the conditions where they were derived. In this work, the cooling rate of centrifugally atomized Al-4%Cu was studied via mathematical modelling and experimental validation. A heat transfer model was implemented, and the value of the convective heat transfer coefficient was obtained from the Whitaker semi-empirical correlation considering three cases of study for the thermophysical properties of the gas. The validity of the Whitaker correlation was experimentally evaluated by means of cooling rates based on the Secondary Dendrite Arm Spacing (SDAS) technique. The Whitaker correlation with the Reynolds and Prandtl numbers evaluated at the ambient temperature and the gas conductivity evaluated at the film temperature gave the best agreement with the experimental results, with cooling rates in the order of 105 Ks−1 for <32.5 µm powders atomized in He atmosphere. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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<p>Schematic diagram of the centrifugal atomization pilot plant.</p>
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<p>Centrifugal atomization process in operation.</p>
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<p>Schematic representation of SDAS measurement.</p>
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<p>Micrographs of Al-4%Cu centrifugal atomized particles (size fraction from 45 to 75 µm) at different gas compositions and melt superheat temperatures: (<b>a</b>) Ar-400 K; (<b>b</b>) Ar-250 K; (<b>c</b>) He-400 K; and (<b>d</b>) He-250 K.</p>
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<p>(<b>a</b>) Experimental results of the SDAS as a function of the mean particle size for a corresponding size range; (<b>b</b>) calculated cooling rate as a function of the mean particle size for a corresponding size range.</p>
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<p>Theoretically calculated cooling rate as a function of the mean particle diameter for particles atomized in Ar and He atmospheres at 250 K superheat temperature. Gas properties were evaluated at three different temperatures: case 1, at gas ambient temperature <span class="html-italic">T<sub>A</sub></span>; case 2, at film temperature <span class="html-italic">T<sub>f</sub></span>; case 3, at droplet surface temperature <span class="html-italic">T<sub>d</sub></span>.</p>
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<p>Comparison between the theoretically calculated cooling rate for the three cases of study and the calculated cooling rates based on SDAS measurements.</p>
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<p>The temperature profile of centrifugally atomized particles for Case 2 of this study, for particles atomized in Ar and He atmospheres at 250 K superheat temperature.</p>
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12 pages, 5171 KiB  
Article
Magnesiothermic Reduction of Silica: A Machine Learning Study
by Kai Tang, Azam Rasouli, Jafar Safarian, Xiang Ma and Gabriella Tranell
Materials 2023, 16(11), 4098; https://doi.org/10.3390/ma16114098 - 31 May 2023
Cited by 3 | Viewed by 1959
Abstract
Fundamental studies have been carried out experimentally and theoretically on the magnesiothermic reduction of silica with different Mg/SiO2 molar ratios (1–4) in the temperature range of 1073 to 1373 K with different reaction times (10–240 min). Due to the kinetic barriers occurring [...] Read more.
Fundamental studies have been carried out experimentally and theoretically on the magnesiothermic reduction of silica with different Mg/SiO2 molar ratios (1–4) in the temperature range of 1073 to 1373 K with different reaction times (10–240 min). Due to the kinetic barriers occurring in metallothermic reductions, the equilibrium relations calculated by the well-known thermochemical software FactSage (version 8.2) and its databanks are not adequate to describe the experimental observations. The unreacted silica core encapsulated by the reduction products can be found in some parts of laboratory samples. However, other parts of samples show that the metallothermic reduction disappears almost completely. Some quartz particles are broken into fine pieces and form many tiny cracks. Magnesium reactants are able to infiltrate the core of silica particles via tiny fracture pathways, thereby enabling the reaction to occur almost completely. The traditional unreacted core model is thus inadequate to represent such complicated reaction schemes. In the present work, an attempt is made to apply a machine learning approach using hybrid datasets in order to describe complex magnesiothermic reductions. In addition to the experimental laboratory data, equilibrium relations calculated by the thermochemical database are also introduced as boundary conditions for the magnesiothermic reductions, assuming a sufficiently long reaction time. The physics-informed Gaussian process machine (GPM) is then developed and used to describe hybrid data, given its advantages when describing small datasets. A composite kernel for the GPM is specifically developed to mitigate the overfitting problems commonly encountered when using generic kernels. Training the physics-informed Gaussian process machine (GPM) with the hybrid dataset results in a regression score of 0.9665. The trained GPM is thus used to predict the effects of Mg-SiO2 mixtures, temperatures, and reaction times on the products of a magnesiothermic reduction, that have not been covered by experiments. Additional experimental validation indicates that the GPM works well for the interpolates of the observations. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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<p>The Mg–SiO<sub>2</sub> pseudo-binary phase diagram calculated by FactSage and its commercial databases [<a href="#B8-materials-16-04098" class="html-bibr">8</a>].</p>
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<p>Calculated (<b>a</b>) iso-Mg and (<b>b</b>) iso-Si content contours in the liquid metal phase.</p>
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<p>Effect of the Mg/SiO<sub>2</sub> ratio on (<b>a</b>) the equilibrium liquid metal phase fraction and (<b>b</b>) total metallic species.</p>
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<p>Schematic drawing of the reactor.</p>
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<p>EPMA mapping of Si, Mg, and O elements in the unreacted core sample.</p>
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<p>EPMA mapping of Si, Mg, and O elements in the reacted core sample.</p>
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<p>Comparison of GPM-calculated species variation as a function of the Mg/SiO<sub>2</sub> molar ratio at 1173 K (<b>a</b>) for 120 min and (<b>b</b>) for 240 min with the experimental results [<a href="#B6-materials-16-04098" class="html-bibr">6</a>,<a href="#B7-materials-16-04098" class="html-bibr">7</a>].</p>
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<p>Comparison of the experimental results [<a href="#B6-materials-16-04098" class="html-bibr">6</a>,<a href="#B7-materials-16-04098" class="html-bibr">7</a>] with GPM-calculated species variations over time for an Mg/SiO<sub>2</sub> molar ratio of 2 at 1173 K.</p>
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<p>(<b>a</b>) Predictive capability of the GPM: phase fractions of species at 1173 K for varying Mg/SiO<sub>2</sub> molar ratios. (<b>b</b>) Comparison of GPM-calculated and measured [<a href="#B6-materials-16-04098" class="html-bibr">6</a>,<a href="#B7-materials-16-04098" class="html-bibr">7</a>] species variations as functions of the reaction time at Mg/SiO<sub>2</sub> ratios of 2 and 2.9 at 1173 K.</p>
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<p>The iso-contours of (<b>a</b>) Mg<sub>2</sub>Si, (<b>b</b>) Si, (<b>c</b>) MgO, and (<b>d</b>) Mg generated by the GPM in comparison with the experimental data points [<a href="#B6-materials-16-04098" class="html-bibr">6</a>,<a href="#B7-materials-16-04098" class="html-bibr">7</a>].</p>
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11 pages, 3452 KiB  
Article
Analyzing the Sintering Kinetics of Ti12.5Ta12.5Nb Alloy Produced by Powder Metallurgy
by Rogelio Macias, Pedro Garnica, Ceylin Fernandez-Salvador, Luis Olmos, Omar Jimenez, Manuel Arroyo-Albiter, Santiago Guevara-Martinez and Jose Luis Cabezas-Villa
Metals 2023, 13(6), 1026; https://doi.org/10.3390/met13061026 - 27 May 2023
Cited by 1 | Viewed by 1308
Abstract
The focus of this work is to analyze the sintering kinetics of Ti12.5Ta12.5Nb alloy by dilatometry. The mixture of powders was achieved by mixing individual powders of Ti, Ta and Nb, which were then axially pressed. Sintering was performed at 1260 °C using [...] Read more.
The focus of this work is to analyze the sintering kinetics of Ti12.5Ta12.5Nb alloy by dilatometry. The mixture of powders was achieved by mixing individual powders of Ti, Ta and Nb, which were then axially pressed. Sintering was performed at 1260 °C using different heating rates. The microstructure was determined by X-ray diffraction and scanning electron microscopy. Results show that densification is achieved by solid state diffusion and that the relative density increased as the heating rate was slow. Due to the full solubility of Ta and Nb in Ti, the relative density reached was up to 93% for all samples. Activation energy was estimated from the densification rate and it was determined that two main diffusion mechanisms were predominant: grain boundary and lattice self-diffusion. This suggests that Ta and Nb diffusion did not affect the atomic diffusion to form the necks between particles. The microstructure shows a combination of α, β and α′, and α″ martensitic phases as a result of the diffusion of Ta and Nb into the Ti unit cell. It was concluded that the heating rate plays a major role in the diffusion of Ta and Nb during sintering, which affects the resulting microstructure. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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<p>Initial powders (<b>a</b>) Ti, (<b>b</b>) Ta and (<b>c</b>) Nb.</p>
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<p>Relative density as a function of temperature during the whole sintering cycle (<b>a</b>), densification rate as a function of the temperature during the heating stage (<b>b</b>) and densification rate as a function of time during isothermal sintering (<b>c</b>).</p>
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<p>Arrhenius plot of ln (<span class="html-italic">T</span> (<span class="html-italic">dl</span>/<span class="html-italic">l</span><sub>0</sub><span class="html-italic">dt</span>)) versus the inverse of absolute temperature (<span class="html-italic">T</span>) at different relative densities.</p>
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<p>X ray patterns of the samples sintered at 1260 °C with different heating rate.</p>
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<p>SEM micrographs of the samples sintered at different heating rates (<b>a</b>) 5 °C/min, (<b>b</b>) 15 °C/min and (<b>c</b>) 25 °C/min. Images close to the remaining particles showing the neighborhood, (<b>d</b>) Nb and (<b>e</b>) Ta particles, respectively. Image a higher magnification showing the microstructure of the β-lamellae (<b>f</b>).</p>
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<p>EDS mapping showing the distribution of the elements on the polished surface of the sample sintered at 5 °C/min.</p>
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<p>EDS mapping showing the martensitic phase of sample sintered at 25 °C/min.</p>
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20 pages, 7933 KiB  
Article
Heat Treatment of AA7075 by Electropulsing and DC Current Application
by Tyler Grimm and Laine Mears
J. Manuf. Mater. Process. 2023, 7(2), 73; https://doi.org/10.3390/jmmp7020073 - 12 Apr 2023
Cited by 3 | Viewed by 2117
Abstract
Electrical resistivity was used in this test methodology to estimate the relative precipitate density in AA7075. Various electrical test parameters were explored to determine the difference between pulsed and DC-type currents. No difference between these test conditions could be distinguished. Furthermore, an electroplastic [...] Read more.
Electrical resistivity was used in this test methodology to estimate the relative precipitate density in AA7075. Various electrical test parameters were explored to determine the difference between pulsed and DC-type currents. No difference between these test conditions could be distinguished. Furthermore, an electroplastic effect was not needed to explain these results and the effects are likely to be caused by purely joule heating. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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<p>Thermal analysis of experiments performed in [<a href="#B3-jmmp-07-00073" class="html-bibr">3</a>]; Legend values indicate convection coefficient (W/(m<sup>2</sup>K)).</p>
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<p>Retrogression hardness of AA7075; Reprinted from Publication [<a href="#B8-jmmp-07-00073" class="html-bibr">8</a>], with permission from Elsevier.</p>
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<p>Boundary conditions used in numerical simulation. Convection cooling also applied on opposite face.</p>
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<p>Example mesh used for numerical simulation.</p>
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<p>Summary result of gage length variation.</p>
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<p>Microstructure of AA7075 sheets after shear cutting; (<b>a</b>) image of several samples, (<b>b</b>–<b>d</b>) close-up images of select samples.</p>
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<p>Resulting resistivity of different heat treatment cycles; Legend values indicate furnace temperature; AQ—air quenched, WQ—water quenched.</p>
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<p>Simulation results of furnace heating and cooling; Legend values represent furnace temperature (<b>a</b>) or initial temperature (<b>b</b>).</p>
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<p>Fixture used to apply current to specimens; (<b>a</b>) front view, (<b>b</b>) back view.</p>
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<p>Current waveform response of Ametek Sorenson SGA 10/1200 power supply.</p>
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<p>Water-cooling, water-quench test condition.</p>
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<p>Current profiles for numerical thermal simulation study; Legend values correspond to tests described in <a href="#jmmp-07-00073-t002" class="html-table">Table 2</a>.</p>
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<p>Temperature results of numerical simulation; Temperatures measured at the center of the geometry.</p>
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<p>Example battery bank output: Added inductance: 1.672 mH, 36 V.</p>
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<p>DC test results of ex situ resistivity measurements; Dashed black lines represent a single standard deviation of the -T6 baseline resistivity; (<b>a</b>) Air-cooled, air-quenched, (<b>b</b>) air-cooled, water-quenched, (<b>c</b>) water-cooled, water-quenched, temperature values are non-physical.</p>
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<p>16 Hz test results of ex situ resistivity measurements, legend values indicate number of pulses applied; Dashed black lines represent a single standard deviation of the -T6 baseline resistivity; (<b>a</b>) Air-cooled, air-quenched, (<b>b</b>) air-cooled, water-quenched.</p>
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<p>Trend line comparison for DC and 16 Hz tests; (<b>a</b>) Air-cooled, air-quenched, (<b>b</b>) air-cooled, water-quenched.</p>
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<p>Results of testing using battery bank as power supply; Legend values correspond to parameters listed in <a href="#jmmp-07-00073-t003" class="html-table">Table 3</a>.</p>
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<p>Current amplitude response of battery bank testing using Test ID 1 with 36 V.</p>
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<p>(<b>a</b>) Trend lines and (<b>b</b>) data points of all tests; Air-cooled, air-quenched.</p>
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<p>Comparison of electrically assisted heat treatment and hardness data presented in [<a href="#B8-jmmp-07-00073" class="html-bibr">8</a>]: circle data points represent [<a href="#B8-jmmp-07-00073" class="html-bibr">8</a>] hardness, star data points represent resistivity values from the present study.</p>
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<p>DC pulse of approximately 200 A/mm<sup>2</sup> using the battery bank power supply.</p>
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<p>DC pulse of approximately 200 A/mm<sup>2</sup> using the Ametek power supply with voltage control; Air-cooled test condition.</p>
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<p>DC pulse of approximately 200 A/mm<sup>2</sup> using the Ametek power supply with voltage control; Water-cooled test condition.</p>
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<p>Direct comparison of testing described in Xu et al. [<a href="#B3-jmmp-07-00073" class="html-bibr">3</a>] and DC pulse testing; Test length: 220 ms.</p>
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<p>Example microstructure of two separate AA7075 sheets.</p>
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<p>Ametek output at 18 Hz; Displays distorted waveshape.</p>
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<p>Trend of percent uncertainty for different maximum temperature values; red points indicate data, blue line indicates fitted trend.</p>
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<p>Attempted trend line fitting to DC test data; (<b>a</b>) Air-cooled, air-quenched, (<b>b</b>) air-cooled, water-quenched, (<b>c</b>) water-cooled, water-quenched, Temperature values are non-physical.</p>
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<p>Example effect of changing voltage on power supply output.</p>
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<p>Output of power supply for all test IDs; 36 V.</p>
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<p>Specimen cooling at various starting temperatures.</p>
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18 pages, 5598 KiB  
Article
Assessment of the Possibility of Reducing Energy Consumption and Environmental Pollution in the Steel Wire Manufacturing Process
by Maciej Suliga, Radosław Wartacz, Joanna Kostrzewa and Marek Hawryluk
Materials 2023, 16(5), 1940; https://doi.org/10.3390/ma16051940 - 26 Feb 2023
Cited by 2 | Viewed by 2378
Abstract
This paper describes research on the influence the technology of zinc-coated steel wire manufacturing has on the energy and force parameters of the drawing process, energy consumption and zinc expenditure. In the theoretical part of the paper, the theoretical work and drawing power [...] Read more.
This paper describes research on the influence the technology of zinc-coated steel wire manufacturing has on the energy and force parameters of the drawing process, energy consumption and zinc expenditure. In the theoretical part of the paper, the theoretical work and drawing power were calculated. Calculations of the electric energy consumption have shown that usage of the optimal wire drawing technology results in a 37% drop in energy consumption, which in the course of a single year translates to savings equal to 13 TJ. This, in turn, results in the decrease of CO2 emissions by tons and a total decrease of the eco-costs by approximately EUR 0.5 mln. Drawing technology also influences the losses of the zinc coating and CO2 emissions. Properly adjusted parameters of the wire drawing technology allow obtaining a zinc-coating that is 100% thicker, translating to 265 tons of zinc, whose production generates 900 tons of CO2 and incurs eco-costs equal to EUR 0.6 mln. Optimal parameters for drawing, from the perspective of decreased CO2 emissions during the zinc-coated steel wire manufacturing, are as follows: usage of the hydrodynamic drawing dies, angle of the die reducing zone α = 5°, and drawing speed of 15 m/s. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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<p>Model and parameters of the zinc-coated steel wire drawing process using a Mario Frigerio multi-stage drawing machine, where α—the angle of the die compression zone, v—drawing speed, d<sub>0</sub>, d<sub>1</sub>—initial and end diameter of the wire.</p>
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<p>Influence of the drawing angle on the total drawing power N<sub>ct</sub> of the wire with an initial diameter of 5.5 mm and end diameter of 2.2 mm for wires drawn with speed v = 5, 10, 15, 20 m/s.</p>
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<p>Influence of the drawing angle α on the theoretical wire drawing power in respective drafts; v = 10 m/s.</p>
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<p>Influence of the drawing speed v on the total drawing power in conventional dies (K) and hydrodynamic dies (H).</p>
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<p>Theoretical drawing power in respective drafts for the wire drawn in conventional dies (K) and hydrodynamic dies (H); v = 10 m/s.</p>
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<p>Power consumption reading field by the Mario Frigerio machine and the total drawing power as a function of a drawing angle for the wires drawn with a speed of v = 5, 10, 15 and 20 m/s.</p>
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<p>Control panel of the Mario Frigerio machine and percentage load of the engines P in respective drafts for the wires drawn with the angle α = 3, 4, 5, 6 and 7° and drawing speed of 20 m/s.</p>
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<p>Total drawing power N<sub>c</sub> as a function of the drawing speed for the wires manufactured using conventional K and hydrodynamic H methods.</p>
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<p>Percentage engine load P in respective drafts for the wires manufactured using conventional K and hydrodynamic H methods.</p>
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<p>Change of the total drawing power, both theoretical and actual, as a function of the drawing angle for the wires drawn with the speed of 20 m/s.</p>
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<p>Change of the total theoretical and actual drawing power for hydrodynamic dies as a function of drawing speed.</p>
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<p>Influence of the technology of wire manufacturing on energy consumption in relation to the manufacturing of 100,000 tons of steel wire drawn with speed v = 5, 10, 15, 20 m/s.</p>
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<p>Influence of the technology of wire manufacturing on the CO<sub>2</sub> emissions caused by energy consumption in relation to the production of 100,000 tons of steel wire drawn with v = speed 5, 10, 15, 20 m/s.</p>
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<p>Influence of the technology of wire manufacturing on the CO<sub>2</sub> emissions caused by energy consumption in relation to the production of 100,000 tons of steel wire drawn with speed v = 5, 10, 15, 20 m/s.</p>
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<p>Influence of the technology of wire manufacturing on the generation of costs (in Euro) caused by energy consumption in relation to the production of 100,000 tons of steel wire.</p>
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<p>Influence of the technology of wire manufacturing on the mass of zinc coating in relation to 100,000 tons of steel wire drawn with speed v = 5, 10, 15, 20 m/s.</p>
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<p>Costs generated by the loss of zinc in the production of 100,000 tons of steel wire.</p>
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17 pages, 12468 KiB  
Article
Evaluation of the Metallurgical Quality of Nodular Cast Iron in the Production Conditions of a Foundry
by Rafał Dwulat and Krzysztof Janerka
J. Manuf. Mater. Process. 2023, 7(1), 18; https://doi.org/10.3390/jmmp7010018 - 4 Jan 2023
Cited by 4 | Viewed by 3005
Abstract
The aim of this research was to determine the factors affecting the metallurgical quality of cast iron during serial production of castings using a campaign cupola and a holding furnace. The problem to be solved, which was to obtain cast iron with the [...] Read more.
The aim of this research was to determine the factors affecting the metallurgical quality of cast iron during serial production of castings using a campaign cupola and a holding furnace. The problem to be solved, which was to obtain cast iron with the required mechanical properties while reducing the internal porosity, results from the foundry’s need to increase the metallurgical quality of the alloy. The increasing difficulty and complicated constructions of castings, for which it is not possible to introduce risers at the stage of technological design, make the stage of proper preparation of cast iron the only way to obtain castings without shrinkage defects. The article presents the results of the study of physicochemical and mechanical properties, microstructure and shrinkage tendency of ductile iron depending on the charge materials used, the amount of Mg used during spheroidization and the type of final inoculants. Step castings and wedge tests were produced on a vertical molding line. The spheroidization was carried out by injecting a core wire containing Mg alloy into the cast iron. The final inoculation of 0.2% was performed using a pneumatic dispenser equipped with a vision system to control the effectiveness of the inoculation. The ITACA Meltdeck thermal analysis system was used to study the physicochemical properties of the initial cast iron, and the ITACA X system to analyze the state of the final cast iron on the molding line. Mechanical tests were performed on samples cut from a stepped casting, and microstructure tests were carried out using a light microscope and a scanning electron microscope. The results of thermal analyses show that increasing the share of pig iron at the expense of steel increases the minimum solidification temperature of eutectic, and thus, increases the potential for graphite nucleation in cast iron. Increasing the nucleation potential can be obtained by adding anthracite, FeSi and SiC. A very important factor in obtaining cast iron of high metallurgical quality is the possible limitation during spheroidization of the length of the core wire containing Mg, which is a carbide-forming element. The lower the initial sulfur level, the greater the possibility of reducing the amount of cored wire. The inoculants containing Ce and Bi were the most advantageous final inoculants from the point of view of obtaining the best microstructure parameters and plastic properties of cast iron. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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<p>Diagram of cast iron melt.</p>
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<p>An example of a thermal analysis curve.</p>
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<p>Experimental pattern plate.</p>
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<p>Fracture of sample inoculated with Zr.</p>
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<p>Fracture of sample inoculated with Ce.</p>
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<p>Fracture of sample inoculated with Ce.</p>
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<p>Fracture of sample inoculated with Ba.</p>
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<p>Fracture of sample inoculated with Bi.</p>
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<p>The obtained Temin values depending on the share of pig iron.</p>
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<p>The obtained Temin values depending on the share of steel scrap.</p>
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<p>The obtained VPS values depending on the share of pig iron.</p>
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<p>The obtained VPS values depending on the share of additional ingredients anthracite + FeSi + SiC.</p>
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<p>The obtained Rec values depending on the share of pig iron.</p>
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<p>The amount of Mg consumed during spheroidization depending on the level of the initial sulfur.</p>
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<p>The obtained Temin values depending on the amount of Mg consumed during spheroidization.</p>
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<p>The obtained VPS values depending on the amount of Mg consumed during magnesium treatment.</p>
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17 pages, 5279 KiB  
Article
Microhardness Distribution of Long Magnesium Block Processed through Powder Metallurgy
by Jiaying Wang and Qizhen Li
J. Manuf. Mater. Process. 2023, 7(1), 5; https://doi.org/10.3390/jmmp7010005 - 27 Dec 2022
Cited by 2 | Viewed by 2563
Abstract
Powder metallurgy is a popular method of making raw powders into specific shaped samples. However, the pressure distribution and the microhardness difference within the sample are nonnegligible and unclear when the sample is long or exceeds a specific size. In this study, the [...] Read more.
Powder metallurgy is a popular method of making raw powders into specific shaped samples. However, the pressure distribution and the microhardness difference within the sample are nonnegligible and unclear when the sample is long or exceeds a specific size. In this study, the long magnesium blocks, with a ratio of about 2.8 between the sample height and the sample side length, are successfully synthesized under three uniaxial and two biaxial conditions. Then, the sample hardness values on the outer surface and the center plane are tested to study the microhardness distribution. The modified analytical expression indicates that the normal pressure exponentially decreases along the compression direction, which is consistent with the hardness distribution trend. Because higher pressure leads to a more compact arrangement of the powders, more metal bonds are formed after sintering. During the first pressing, the sidewall pressure makes the surface hardness higher. The secondary reverse compression mainly improves the bottom and core hardness due to the re-orientation and re-location of the powders. The obtained relationship between the applied pressure and the hardness distribution is instructive in predicting and improving the sample quality. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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Graphical abstract

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<p>(<b>a</b>) Scanning electron microscopic image of pure Mg powders; (<b>b</b>) image of a sintered sample.</p>
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<p>Schematic diagram of sample preparation.</p>
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<p>(<b>a</b>) Geometric diagram of the final sample and sample cut; (<b>b</b>) schematic diagram of force analysis on a slice in pressing die; (<b>c</b>) image of the sample after hardness test.</p>
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<p>Microstructures of uniaxial PM Mg under (<b>a</b>,<b>b</b>) S1, (<b>c</b>,<b>d</b>) S2, and (<b>e</b>,<b>f</b>) S3 conditions.</p>
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<p>Microstructures of biaxial PM Mg under (<b>a</b>,<b>b</b>) D1 and (<b>c</b>,<b>d</b>) D2 conditions.</p>
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<p>Hardness test results of S1−1 sample on the (<b>a</b>) surface and (<b>b</b>) center; hardness contour plot of S1−1 sample (<b>c</b>) surface and (<b>d</b>) center; hardness 3D plot of S1−1 sample (<b>e</b>) surface and (<b>f</b>) center.</p>
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<p>Hardness results of D1−1 sample (<b>a</b>) surface and (<b>b</b>) center; hardness contour plot of D1−1 sample (<b>c</b>) surface and (<b>d</b>) center; hardness 3D plot of D1−1 sample (<b>e</b>) surface and (<b>f</b>) center.</p>
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<p>S1−1, S1−2, and S1−3 sample microhardness scatter plots and S1 microhardness fitting curves on the (<b>a</b>) surface and (<b>b</b>) center; S2−1, S2−2, and S2−3 sample microhardness scatters and S2 microhardness fitting curves on (<b>c</b>) surface and (<b>d</b>) center; S3 sample microhardness scatters and fitting curves, S1 hardness fitting curve, and S2 hardness fitting curve on the (<b>e</b>) surface and (<b>f</b>) center.</p>
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<p>D1−1, D1−2, and D1−3 sample microhardness scatters and fitting curves on (<b>a</b>) surface and (<b>b</b>) center; D2−1, D2−2, and D2−3 sample microhardness scatters and fitting curves on (<b>c</b>) surface and (<b>d</b>) center; D1 and D2 hardness fitting curves on (<b>e</b>) surface and (<b>f</b>) center.</p>
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13 pages, 3376 KiB  
Article
Fuel Consumption Dependence on a Share of Reduction Processes in Imperial Smelting Furnace
by Mikolaj Bernasowski, Ryszard Stachura and Arkadiusz Klimczyk
Energies 2022, 15(23), 9259; https://doi.org/10.3390/en15239259 - 6 Dec 2022
Cited by 3 | Viewed by 2192
Abstract
The paper shows the use of novel modelling techniques adapted from ironmaking in the pyrometallurgical process of zinc production. Firstly, regarding the purpose to determine the boundary conditions of reduction processes taking part in the working volume of an Imperial Smelting Furnace (ISF), [...] Read more.
The paper shows the use of novel modelling techniques adapted from ironmaking in the pyrometallurgical process of zinc production. Firstly, regarding the purpose to determine the boundary conditions of reduction processes taking part in the working volume of an Imperial Smelting Furnace (ISF), a deep thermochemical analysis was conducted. On this basis and using Ramm’s principles of direct and indirect reduction optimal share, the fuel rate minimization model was built. The model’s leading role is minimizing coke consumption in the ISF while maintaining the thermal state of the furnace at the correct level. In addition, the proposed presentation of the ISF thermal state shows in a unified way all the shortcomings in the correct process operation. Verification in real conditions on the ISF in Miasteczko Śląskie shows that model implementation can bring tangible benefits. Coke savings can reach over 30 kg per tonne of raw zinc. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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<p>The ISF working diagram.</p>
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<p>Free Gibbs energy in the temperature dependence for the main reaction in ISP.</p>
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<p>Equilibrium of all stable phases dependent upon the temperature.</p>
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<p>Equilibrium of gas phases dependent upon the temperature.</p>
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<p>Gas-phase-reducing ability dependent upon the temperature for ZnO-CO-C.</p>
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<p>Carbon rate–ZnO reduction ratio diagram.</p>
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<p>Material streams in an ISF.</p>
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<p>View of the fuel rate minimization model working on real data.</p>
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<p>Operation point adjustment by fuel rate minimization model in a one-hour cycle.</p>
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11 pages, 2117 KiB  
Article
Localized Defects in Cold Die-Compacted Metal Powders
by Elisa Torresani, Gloria Ischia and Alberto Molinari
J. Manuf. Mater. Process. 2022, 6(6), 155; https://doi.org/10.3390/jmmp6060155 - 6 Dec 2022
Cited by 2 | Viewed by 3062
Abstract
In powder metallurgy (PM), the compaction step is fundamental to determining the final properties of the sintered components. The deformation and defectiveness introduced in the powder material during uniaxial die compaction can be correlated to the activation and enhancement of the dislocation pipe [...] Read more.
In powder metallurgy (PM), the compaction step is fundamental to determining the final properties of the sintered components. The deformation and defectiveness introduced in the powder material during uniaxial die compaction can be correlated to the activation and enhancement of the dislocation pipe diffusion, a lattice diffusion mechanism during the sintering process. Its coefficient depends on the dislocation density. The powder particles are mostly deformed along the direction of the compaction (longitudinal direction) rather than along the compaction plane; consequently, the contact areas perpendicular to the direction of the compaction present a higher density of dislocations and lattice defects. This high density intensifies the shrinkage along the direction of compaction. To demonstrate the influence of uniaxial cold compaction on the material’s stress state the powder particles and their contacts were modeled using spheres made of pure copper. These spheres are compacted in a die at different pressures to better analyze the system’s response at the grade of deformation and the consequent influence on the material’s behavior during the sintering. In the different zones of the sphere, the micro-hardness was measured and correlated to the concentration of dislocations using the model for indentation size effect (ISE). After the compaction, the spheres were more deformed along the longitudinal than the transversal direction. The results obtained using hardness indentation show differences in the dislocation density between the undeformed and deformed spheres and, in the case of the compacted sphere, between the contact area along the longitudinal and the transversal direction. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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<p>(<b>a</b>) Compaction system involved in the present work, and (<b>b</b>) die dimensions.</p>
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<p>Barycentric section of (<b>a</b>) sphere before the cold-compaction. (<b>b</b>) deformed sphere at 196 MPa and (<b>c</b>) at 500 MPa; the arrows indicate the direction of the compaction. In the red dotted circles, the micro-indentation imprints are shown.</p>
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<p>TEM images obtained (<b>a</b>) on the surface of undeformed spheres, (<b>b</b>) in the contact region.</p>
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<p>SEM images of the sphere compacted at 196.5 MPa: (<b>a</b>) etching pits in contact zone along the direction of compaction and (<b>b</b>) in the transversal plane at 10,000×; (<b>c</b>) Count of pits for the direction parallel to the pressure at 10,000× and (<b>d</b>) transversal at 5000×.</p>
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<p>Data of micro-indentation measurements for the undeformed spheres at the (<b>a</b>) surface and (<b>b</b>) core; compacted spheres at 196 MPa at the (<b>c</b>) perpendicular contacts, (<b>d</b>) parallel contacts and (<b>e</b>) core; compacted spheres at 500 MPa at the (<b>f</b>) perpendicular contacts, (<b>g</b>) parallel contacts and (<b>h</b>) core.</p>
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11 pages, 2562 KiB  
Article
Investigation of the OA-300M Electrolysis Cell Temperature Field of Metallurgical Production
by Tatyana Valeryevna Kukharova, Yury Valeryevich Ilyushin and Mir-Amal Mirrashidovich Asadulagi
Energies 2022, 15(23), 9001; https://doi.org/10.3390/en15239001 - 28 Nov 2022
Cited by 40 | Viewed by 1983
Abstract
In this study, the authors explored the issues of the Soderbergh electrolysis cell’s increase in economic efficiency. This result was achieved by analyzing the temperature fields of the electrolysis cell in order to determine the overheating points. As a result, it led to [...] Read more.
In this study, the authors explored the issues of the Soderbergh electrolysis cell’s increase in economic efficiency. This result was achieved by analyzing the temperature fields of the electrolysis cell in order to determine the overheating points. As a result, it led to the determination of the points of the hearth bottom’s subsequent breakdown, causing the failure of the electrolysis cell. In this paper, the mathematical modeling of the temperature fields using a spatially distributed mathematical model and conducted experimental studies were carried out. The mathematical model also provides the opportunity to measure the temperature field in the hearth bottom (at the bottom) of the OA-300M electrolysis cell. The results of the given research can be used to solve the experimental determination of the hearth bottom internal defect problem. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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<p>The object in question.</p>
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<p>First heating pulse (the temperature and scale of the object are indicated in conventional units).</p>
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<p>The system is in the process of reaching a stable state (the temperature and scale of the object are indicated in conventional units).</p>
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<p>Longitudinal section of the OA-300M electrolysis cell [<a href="#B31-energies-15-09001" class="html-bibr">31</a>].</p>
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<p>Overheating of cathode No. 5 and part of the electrolysis cell bottom: (<b>a</b>) overheating area 4–5; (<b>b</b>) overheating area 5–6.</p>
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<p>Overheating of the cathode part of the electrolysis cell bottom: (<b>a</b>) Hearth bottom with pronounced overheating of the right side (cathode 12); (<b>b</b>) Hearth bottom with pronounced overheating of both cathodes (cathodes 17–18).</p>
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<p>Overheating of the corner part of the hearth bottom at different phases of operation: (<b>a</b>) Overheating area at the initial phase of aluminum production; (<b>b</b>) Overheating area at the final phase of aluminum production.</p>
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8 pages, 1746 KiB  
Article
Inclusion Characteristics in Steel with CeO2 Nanoparticle Addition
by Hui Kong, Xiang Cheng, Shoulin Huang and Yue Qiu
Metals 2022, 12(11), 1994; https://doi.org/10.3390/met12111994 - 21 Nov 2022
Cited by 3 | Viewed by 1570
Abstract
The application of Ce oxides in oxide metallurgy has received extensive attention, but until now, the direct adding of CeO2 into molten steel to generate Ce oxides has not occurred. In this paper, a mixture of CeO2 and Si nanoparticles were [...] Read more.
The application of Ce oxides in oxide metallurgy has received extensive attention, but until now, the direct adding of CeO2 into molten steel to generate Ce oxides has not occurred. In this paper, a mixture of CeO2 and Si nanoparticles were added into molten steel. The resultant formation of micrometer scale Ce-bearing oxides confirmed its adding validity. This behavior may be interpreted as the reactivity between CeO2 and [Al], and the improved wettability between CeO2 and molten steel with the assistance of Si powder. Thus, when the quantity of CeO2 is kept constant, its added yield should increase when increasing the added quantity of Si. This was verified by the larger percentage of Ce-bearing oxides of the total oxides and the greater average content of Ce in Ce-bearing oxides after normalization. Moreover, compared with the blank sample, statistical results indicated that the oxides in CeO2-modified samples were refined, and their dispersion homogeneity was enhanced. This comparison indicates the effectiveness of the external adding method in oxide metallurgy. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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<p>SEM micrograph and EDS mapping images of various elements for typical oxide in 1#.</p>
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<p>SEM micrograph and EDS mapping images of various elements for typical oxide in 2#.</p>
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<p>SEM micrograph and EDS mapping images of various elements for typical oxide in 3#.</p>
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<p>The size distribution of oxides for samples.</p>
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13 pages, 2963 KiB  
Article
Analysis of Using Soot Application in the Processing of Zinc-Bearing Waste Materials
by Albert Smalcerz, Szymon Ptak, Jerzy Łabaj, Marzena Półka, Adam Kula and Leszek Blacha
Energies 2022, 15(21), 7969; https://doi.org/10.3390/en15217969 - 27 Oct 2022
Cited by 1 | Viewed by 1682
Abstract
In metallurgical processes, coke is used, among other materials, in order to implement the process of removing zinc from waste by reduction and evaporation. Due to the implemented de-carbonization policy, we are dealing with an increase in costs and decreasing availability of coke, [...] Read more.
In metallurgical processes, coke is used, among other materials, in order to implement the process of removing zinc from waste by reduction and evaporation. Due to the implemented de-carbonization policy, we are dealing with an increase in costs and decreasing availability of coke, which leads to an intensive search for other possibilities for conducting the process, which may generate a fire and explosion hazard in the technological process. This article analyzes the possibility of using soot in the process of reducing the zinc content in deposited metallurgical waste, taking into account the issue of fire and explosion safety. The results of the research proved the possibility of the safe use of the reductor, which is soot and anthracite, as a material replacing coke in pyrometallurgical processes. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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<p>Scheme of the Weltz process with rotary kiln [<a href="#B12-energies-15-07969" class="html-bibr">12</a>,<a href="#B13-energies-15-07969" class="html-bibr">13</a>]: the colors of the drawing correspond to the zones of the furnace (yellow, orange, and red zones are 1, 2, and 3, respectively).</p>
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<p>Scheme of the rotary kiln divided into zones [<a href="#B12-energies-15-07969" class="html-bibr">12</a>]. A detailed description of the reactions in each zone is provided below.</p>
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<p>Diagram of the granulate preparation system used in the research: 1—input material feeding system; 2—intensive mixer; 3—plate mixer; 4—granules.</p>
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<p>Laser diffraction particle size analyzer Fritsch Analysette 22.</p>
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<p>Analyzer NETZSCH STA 449 Jupiter F1 coupled with QMS Aëolos analyzer.</p>
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<p>Twenty-liter spherical vessel manufactured by ANKO S.A.</p>
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<p>Diagram of the laboratory furnace used in the research: 1—Natural gas and air supply; 2—burner; 3—rotary kiln; 4—furnace drive; 5—rollers enabling rotary movement; 6—chute hoppers (return, burnouts); 7—process gas collector; 8—process gas cooler; 9—pulse fabric filter; 10—fan with a chimney system.</p>
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<p>View of the laboratory rotary kiln.</p>
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<p>Particle size distribution of soot dust sample.</p>
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<p>The results of the TG analysis of the soot dust sample.</p>
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13 pages, 6612 KiB  
Article
Fusion Separation of Vanadium-Titanium Magnetite and Enrichment Test of Ti Element in Slag
by Shuangping Yang, Shouman Liu, Shijie Guo, Tiantian Zhang and Jianghan Li
Materials 2022, 15(19), 6795; https://doi.org/10.3390/ma15196795 - 30 Sep 2022
Cited by 4 | Viewed by 2076
Abstract
In view of the problem that the enrichment and migration law of the Ti element in the slag of vanadium-titanium magnetite during the melting process is not clear, the phase transformation is not clear and the enrichment effect is not obvious, the single [...] Read more.
In view of the problem that the enrichment and migration law of the Ti element in the slag of vanadium-titanium magnetite during the melting process is not clear, the phase transformation is not clear and the enrichment effect is not obvious, the single factor experiment and orthogonal experiment are used to optimize the melting conditions of Ti enrichment. Through XRD, SEM and EDS analysis, the effects of melting temperature, alkalinity and carbon content on the Ti phase in the slag are studied, and the occurrence form and migration law of the Ti element in the slag system under different melting conditions are clarified. The results demonstrate that increasing the basicity and melting temperature is beneficial to the enrichment of Ti, but it is too high it will lead to the formation of pyroxene, diopside and magnesia-alumina spinel, affecting the enrichment of Ti. The increase in carbon content can make Ti occur in slag in the form of titanium oxides such as TiO, TiO2, Ti2O3 and Ti3O5, but excessive carbon content leads to the excessive reduction of Ti compounds to TiCN and TiC. After optimization, under the melting conditions of alkalinity 1.2, the melting temperature 1500 °C and carbon content 15%, the content of Ti in slag can reach 18.84%, and the recovery rate is 93.54%. By detecting the content of Fe and V in molten iron, the recovery rates are 99.86% and 95.64%, respectively. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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<p>Particle size distribution of vanadium titanium magnetite.</p>
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<p>XRD patterns of vanadium titanium magnetite.</p>
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<p>Experimental flow chart.</p>
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<p>XRD analysis of metalized pellets.</p>
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<p>Effect of alkalinity change on Ti/Fe content in slag.</p>
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<p>Effect of alkalinity on slag phase XRD. (<b>a</b>) Alkalinity = 1.0; (<b>b</b>) Alkalinity = 1.1; (<b>c</b>) Alkalinity = 1.2; (<b>d</b>) Alkalinity = 1.3.</p>
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<p>Effect of melting temperature on Ti/Fe content in slag.</p>
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<p>Effect of melting temperature on slag phase XRD. (<b>a</b>) 1450 ℃; (<b>b</b>) 1480 ℃; (<b>c</b>) 1500 ℃; (<b>d</b>) 1520 ℃; (<b>e</b>) 1550 ℃.</p>
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<p>Effect of carbon content on Ti/Fe content in slag.</p>
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<p>Effect of carbon content on XRD of slag phase. (<b>a</b>) Carbon content—5%; (<b>b</b>) Carbon content—10%; (<b>c</b>) Carbon content—15%; (<b>d</b>) Carbon content—20%.</p>
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<p>SEM and EDS analysis of slag phase. (<b>a</b>) SEM; (<b>b</b>) EDS.</p>
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13 pages, 20829 KiB  
Article
Effect of Dynamic Preheating on the Thermal Behavior and Mechanical Properties of Laser-Welded Joints
by Linyi Xie, Wenqing Shi, Teng Wu, Meimei Gong, Detao Cai, Shanguo Han and Kuanfang He
Materials 2022, 15(17), 6159; https://doi.org/10.3390/ma15176159 - 5 Sep 2022
Cited by 4 | Viewed by 2274
Abstract
The high cooling rate and temperature gradient caused by the rapid heating and cooling characteristics of laser welding (LW) leads to excessive thermal stress and even cracks in welded joints. In order to solve these problems, a dynamic preheating method that uses hybrid [...] Read more.
The high cooling rate and temperature gradient caused by the rapid heating and cooling characteristics of laser welding (LW) leads to excessive thermal stress and even cracks in welded joints. In order to solve these problems, a dynamic preheating method that uses hybrid laser arc welding to add an auxiliary heat source (arc) to LW was proposed. The finite element model was deployed to investigate the effect of dynamic preheating on the thermal behavior of LW. The accuracy of the heat transfer model was verified experimentally. Hardness and tensile testing of the welded joint were conducted. The results show that using the appropriate current leads to a significantly reduced cooling rate and temperature gradient, which are conducive to improving the hardness and mechanical properties of welded joints. The yield strength of welded joints with a 20 A current for dynamic preheating is increased from 477.0 to 564.3 MPa compared with that of LW. Therefore, the use of dynamic preheating to reduce the temperature gradient is helpful in reducing thermal stress and improving the tensile properties of the joint. These results can provide new ideas for welding processes. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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<p>HLAW system.</p>
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<p>Tensile sample showing dimensions (in millimeters).</p>
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<p>FEM simulation framework.</p>
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<p>Gauss heat source model: (<b>a</b>) LW; (<b>b</b>) HLAW.</p>
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<p>Mesh of FEM.</p>
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<p>Temperature distribution of the arc with currents of (<b>a</b>) 20 A and (<b>b</b>) 40 A.</p>
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<p>Comparison of experimental temperature verification. (<b>a</b>) Temperature curve at point A (30,30,1). (<b>b</b>) Comparison of cross-section molten pool morphology with a 0 A current. (<b>c</b>) Comparison of cross-section molten pool morphology with a 20 A current.</p>
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<p>Temperature distribution at 2.5 s with currents of (<b>a</b>) 0 A and (<b>b</b>) 20 A. (<b>c</b>,<b>d</b>) Magnified views of the central region taken from the central areas in (<b>a</b>,<b>b</b>), respectively.</p>
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<p>Temperature distribution with different currents at special points. (<b>a</b>) Maximum temperature gradients at special points (50, 0, 1). (<b>b</b>) Maximum cooling rate at special points (50, 0, 1). (<b>c</b>) Maximum temperature gradients at special points (<span class="html-italic">x</span>, 0, 1). (<b>d</b>) Maximum cooling rate at special points (<span class="html-italic">x</span>, 0, 1).</p>
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<p>Microhardness distribution in the weld section. (<b>a</b>) Microhardness test position. (<b>b</b>) Microhardness test results.</p>
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<p>Microstructural images obtained from the optical microscope for the WMs. (<b>a</b>–<b>e</b>) 0 A, 10 A, 20 A, 30 A, and 40 A currents, respectively.</p>
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<p>Tensile mechanical properties.</p>
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<p>Scanning electron microscope images of the fracture for (<b>a</b>) 20 A and (<b>b</b>) 40 A cases. (<b>c</b>) and (<b>d</b>) Higher magnification micrographs taken from the central areas in (<b>a</b>) and (<b>b</b>), respectively.</p>
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17 pages, 13414 KiB  
Article
Microstructural Evolution of the TLP Joints of RAFM Steel during Aging and Creep
by Kun Liu, Wenchao Li, Ran Ding and Chenxi Liu
Metals 2022, 12(8), 1333; https://doi.org/10.3390/met12081333 - 10 Aug 2022
Cited by 1 | Viewed by 2312
Abstract
In this study, transient liquid-phase (TLP) bonding was adopted to obtain a reliable reduced-activation ferritic/martensitic (RAFM) steel joint with Fe-Si-B amorphous foil. The aging tests and creep tests of the TLP joints were carried out at 550 °C to study the microstructural evolution [...] Read more.
In this study, transient liquid-phase (TLP) bonding was adopted to obtain a reliable reduced-activation ferritic/martensitic (RAFM) steel joint with Fe-Si-B amorphous foil. The aging tests and creep tests of the TLP joints were carried out at 550 °C to study the microstructural evolution in the service process. The effect of stress loading on the microstructural evolution of the TLP joint was investigated. The results show that creep fractures in the TLP joints occur in the base material. The main factors affecting the creep performance of TLP joints are the recovery of substructures and the coarsening and deformation of martensitic laths. In addition, the M23C6 carbides in the base material were coarser than in the weld zone. Compared with aging samples and creep samples undergoing the same test temperature, the dislocation density in the isothermal solidification zone (ISZ) increased significantly with increases in the stress level. Furthermore, it is worth noting that the microstructure of the weld zone changed from large-sized ferrite to a mixed, fine microstructure of ferrite and martensite, which increases the heat resistance of the TLP joints, and thus results in creep fractures in the base metal. Full article
(This article belongs to the Topic Advanced Processes in Metallurgical Technologies)
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<p>Schematic of the creep specimen (Unit: mm).</p>
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<p>Microstructures of TLP joints after aging at 550 °C with different aging times: (<b>a</b>) 5 h, (<b>b</b>) 20 h, (<b>c</b>) 100 h, and (<b>d</b>) 500 h.</p>
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<p>Grain-boundary-distribution maps of joint regions for different aging times: (<b>a</b>) 5 h, (<b>b</b>) 20 h, (<b>c</b>) 100 h, and (<b>d</b>) 500 h. (The blue lines represent the high-angle boundaries, while the red lines denote the low-angle boundaries.)</p>
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<p>Statistical diagrams of the grain-size of the joint regions for different aging times: (<b>a</b>) 5 h, (<b>b</b>) 20 h, (<b>c</b>) 100 h, and (<b>d</b>) 500 h.</p>
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<p>Creep curves of TLP joints for different stresses: (<b>a</b>) relationship between creep strain and time; (<b>b</b>) relationship been creep rate and time.</p>
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<p>Fracture morphology of the creep-fractured samples for the different stress levels: (<b>a</b>) 160 MPa, (<b>b</b>) 180 MPa, and (<b>c</b>) 220 MPa.</p>
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<p>Microstructure of the sample after creep rupture for the stress of 160 MPa near the fracture. (The yellow arrow denotes the direction of the applied load.)</p>
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<p>Microstructure of the joint area after creep fracture for the different stresses of (<b>a</b>) 160 MPa, (<b>b</b>) 180 MPa, and (<b>c</b>) 220 MPa. (The yellow arrow denotes the direction of the applied load.)</p>
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<p>SEM images of the TLP joint after creep fracture under a stress of 220 MPa: (<b>a</b>) the low-magnification image, (<b>b</b>) the distribution of the M<sub>23</sub>C<sub>6</sub> carbides, and (<b>c</b>) the EDS analysis of the M<sub>23</sub>C<sub>6</sub> carbide indicated by circle in (<b>b</b>).</p>
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<p>TEM images of the weld zone and base metal after creep-fracture for the different stress levels: (<b>a1</b>) base metal, 160 MPa; (<b>a2</b>) weld zone, 160 MPa; (<b>b1</b>) base metal, 180 MPa; (<b>b2</b>) weld zone, 180 MPa; (<b>c1</b>) base metal, 220 MPa; and (<b>c2</b>) weld zone, 220 MPa.</p>
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<p>TEM images of the weld zone and base metal after creep-fracture for the different stress levels: (<b>a1</b>) base metal, 160 MPa; (<b>a2</b>) weld zone, 160 MPa; (<b>b1</b>) base metal, 180 MPa; (<b>b2</b>) weld zone, 180 MPa; (<b>c1</b>) base metal, 220 MPa; and (<b>c2</b>) weld zone, 220 MPa.</p>
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<p>The EBSD results of the weld zone of the creep samples for different stresses: (<b>a</b>–<b>c</b>) are IPF maps; (<b>d</b>–<b>f</b>) are grain-size statistical diagrams.</p>
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<p>The EBSD results of the weld zone of the creep samples for different stresses: (<b>a</b>–<b>c</b>) are grain-boundary distribution maps; (<b>d</b>–<b>f</b>) are orientation-difference statistics diagrams.</p>
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<p>KAM maps of the joint regions for different aging times: (<b>a</b>) 5 h, (<b>b</b>) 20 h, (<b>c</b>) 100 h, and (<b>d</b>) 500 h.</p>
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<p>KAM maps of the weld zone of creep samples for different stresses: (<b>a</b>) 3-550-160, (<b>b</b>) 3-550-180, and (<b>c</b>) 3-550-220.</p>
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<p>Microstructure of the base-material region of creep samples for different stresses: (<b>a</b>) 3-550-160; (<b>b</b>) 3-550-180; and (<b>c</b>) 3-550-220.</p>
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<p>GND maps of the joint regions for different aging times: (<b>a</b>) 5 h, (<b>b</b>) 20 h, (<b>c</b>) 100 h, and (<b>d</b>) 500 h.</p>
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<p>GND maps of the weld zone of creep samples for different stresses: (<b>a</b>) 3-550-160, (<b>b</b>) 3-550-180, and (<b>c</b>) 3-550-220.</p>
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