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J. Manuf. Mater. Process., Volume 5, Issue 1 (March 2021) – 26 articles

Cover Story (view full-size image): The manufacture of thin walls with sharp corners has been optimized by adjusting the limits of a 3-axis Cartesian kinematics through data recorded and analyzed offline, such as axis speed, acceleration, and the positioning of the X and Y axes. The study was carried out with SS316L and IN718 alloys using the directed energy deposition process with laser. Thin walls were obtained with 1 mm thickness, and only one bead per layer and straight/sharp corners at 90. After adjusting the in-position parameter G502 for positioning precision on the FAGOR 8070 CNC system, it was possible to obtain walls with minimal material accumulation in the corner, and practically constant layer thickness and height, with adequate radii of internal curvature. View this paper.
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14 pages, 13111 KiB  
Article
Condition Monitoring of Manufacturing Processes under Low Sampling Rate
by Gabriel Bernard, Sofiane Achiche, Sébastien Girard and René Mayer
J. Manuf. Mater. Process. 2021, 5(1), 26; https://doi.org/10.3390/jmmp5010026 - 23 Mar 2021
Cited by 3 | Viewed by 2622
Abstract
Manufacturing processes can be monitored for anomalies and failures just like machines, in condition monitoring and prognostic and health management. This research takes inspiration from condition monitoring and prognostic and health management techniques to develop a method for part production process monitoring. The [...] Read more.
Manufacturing processes can be monitored for anomalies and failures just like machines, in condition monitoring and prognostic and health management. This research takes inspiration from condition monitoring and prognostic and health management techniques to develop a method for part production process monitoring. The contribution brought by this paper is an automated technique for process monitoring that works with low sampling rates of 1/3Hz, a limitation that comes from using data provided by an industrial partner and acquired from industrial manufacturing processes. The technique uses kernel density estimation functions on machine tools spindle load historical time signals for distribution estimation. It then uses this estimation to monitor the manufacturing processes for anomalies in real time. A modified version was tested by our industrial partner on a titanium part manufacturing line. Full article
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Graphical abstract

Graphical abstract
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<p>Diagram of data relationship at our industrial partner.</p>
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<p>High level diagram of a machining program for complex parts manufacturing used by our industrial partner.</p>
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<p>Spindle load through time compared with historical data.</p>
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<p>Raw data for the program 1543 and tool 6040 on the T5-1 machine.</p>
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<p>Raw data for the program 1543 and tool 6040 on the T5-1 machine for run 127.</p>
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<p>Spindle load probability density estimation on time bins of 3 s for program 1543, tool 6040 and machine T5-1 represented as a heat map.</p>
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<p>Program 1543 tool 6040 on machine T5-1 normal behavior extraction.</p>
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<p>Normal behavior without run 237 and 241 and raw data on top.</p>
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<p>Testing the normal behavior monitoring of program behavior.</p>
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20 pages, 26949 KiB  
Article
Experimental Study on Joining by Forming of HCT590X + Z and EN-AW 6014 Sheets Using Cold Extruded Pin Structures
by David Römisch, Martin Kraus and Marion Merklein
J. Manuf. Mater. Process. 2021, 5(1), 25; https://doi.org/10.3390/jmmp5010025 - 17 Mar 2021
Cited by 11 | Viewed by 3161
Abstract
Due to stricter emission targets in the mobility sector and the resulting trend towards lightweight construction in order to reduce weight and consequently emissions, multi-material systems that allow a material to be placed in the right quantity and in the right place are [...] Read more.
Due to stricter emission targets in the mobility sector and the resulting trend towards lightweight construction in order to reduce weight and consequently emissions, multi-material systems that allow a material to be placed in the right quantity and in the right place are becoming increasingly important. One major challenge that is holding back the rapid and widespread use of multi-material systems is the lack of adequate joining processes that are suitable for joining dissimilar materials. Joining processes without auxiliary elements have the advantage of a reduced assembly effort and no additional added weight. Conventional joining processes without auxiliary elements, such as welding, clinching, or the use of adhesives, reach their limits due to different mechanical properties and chemical incompatibilities. A process with potential in the field of joining dissimilar materials is joining without an auxiliary element using pin structures. However, current pin manufacturing processes are mostly time-consuming or can only be integrated barely into existing industrial manufacturing processes due to their specific properties. For this reason, the present work investigates the production of single- and multi-pin structures from high-strength dual-phase steel HCT590X + Z (DP600, t0 = 1.5 mm) by cold extrusion directly out of the sheet metal. These structures are subsequently joined with an aluminium sheet (EN AW-6014-T4, t0 = 1.5 mm) by direct pin pressing. For a quantitative evaluation of the joint quality, tensile shear tests are carried out and the influence of different pin heights, pin number, and pin arrangements, as well as different joining strategies on the joint strength is experimentally evaluated. It is proven that a single pin structure with a diameter of 1.5 mm and an average height of 1.86 mm achieves a maximum tensile shear force of 1025 N. The results reveal that the formation of a form-fit during direct pin pressing is essential for the joint strength. By increasing the number of pins, a linear increase in force could be demonstrated, which is independent of the arrangement of the pin structures. Full article
(This article belongs to the Special Issue Metal Forming and Joining)
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Figure 1
<p>Process schematic for extruding pin structures from the sheet metal plane.</p>
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<p>Schematic illustration of the investigated pin arrangements.</p>
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<p>Schematic illustration of direct pin pressing with a die.</p>
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<p>Schematic illustration of tensile shear specimen and test setup.</p>
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<p>Nonlinear correlation of the measured pin height to the punch penetration depth.</p>
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<p>Force-displacement curves in relation to the number and arrangement of the pins.</p>
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<p>Illustration of the surface profile of a single and multi-pin structure as well as the sheet thickness across the width of the sample. (<b>a</b>) Surface profiles of the measured specimens, (<b>b</b>) corresponding sheet thickness, (<b>c</b>) detailed view of the red area marked in subfigure (<b>b</b>) and (<b>d</b>) detailed view of the blue area marked in subfigure (<b>b</b>).</p>
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<p>Force-displacement curves of the direct pin pressing joining process. (<b>a</b>) Force-displacement curves when joining 1.08 mm pins without and with die, (<b>b</b>) Force-displacement curves when joining 1.45 mm pins without and with die, (<b>c</b>) Force-displacement curves when joining 1.86 mm pins without and with die and (<b>d</b>) Force-displacement curves when joining 1.55 mm multi pin structures.</p>
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<p>Comparison of the force displacement curves of the two investigated joining strategies based on the 1.86 mm pin structures.</p>
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<p>Micrographs of the investigated joint variations. (<b>a</b>) 1.08 mm pin joined without die, (<b>b</b>) 1.08 mm pin joined with 3 mm die, (<b>c</b>) 1.08 mm pin joined with 4 mm die, (<b>d</b>) 1.45 mm pin joined without die, (<b>e</b>) 1.45 mm pin joined with 3 mm die, (<b>f</b>) 1.45 mm pin joined with 4 mm die, (<b>g</b>) 1.86 mm pin joined without die, (<b>h</b>) 1.86 mm pin joined with 3 mm die, (<b>i</b>) 1.86 mm pin joined with 4 mm die, (<b>j</b>) 1.55 mm longitudinal multi pins and (<b>k</b>) 1.55 mm transverse multi pins.</p>
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<p>Tensile shear force–displacement curves of the investigated pin-height/joining strategy combination as well as multi pin joints. (<b>a</b>) Tensile shear force-displacement curves with corresponding maximum force for 1.08 mm single pins joined without and with die, (<b>b</b>) tensile shear force-displacement curves with corresponding maximum force for 1.45 mm single pins joined without and with die, (<b>c</b>) tensile shear force-displacement curves with corresponding maximum force for 1.86 mm single pins joined without and with die and (<b>d</b>) tensile shear force-displacement curves with corresponding maximum force for 1.55 mm multi pins.</p>
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<p>Damage patterns of the examined joints with steel sheet containing the pin structure at the top and corresponding aluminium sheet at the bottom. (<b>a_1</b>) 1.08 mm pin joined without die, (<b>a_2</b>) 1.08 mm pin joined with 3 mm die, (<b>a_3</b>) 1.08 mm pin joined with 4 mm die, (<b>b_1</b>) 1.45 mm pin joined without die, (<b>b_2</b>) 1.45 mm pin joined with 3 mm die, (<b>b_3</b>) 1.45 mm pin joined with 4 mm die, (<b>c_1</b>) 1.86 mm pin joined without die, (<b>c_2</b>) 1.86 mm pin joined with 3 mm die, (<b>c_3</b>) 1.86 mm pin joined with 4 mm die, (<b>d</b>) 1.55 mm longitudinal multi pins and (<b>e</b>) 1.55 mm transverse multi pins.</p>
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2 pages, 164 KiB  
Editorial
Cyber-Physical Production Systems (CPPS): Introduction
by Sebastian Thiede
J. Manuf. Mater. Process. 2021, 5(1), 24; https://doi.org/10.3390/jmmp5010024 - 17 Mar 2021
Cited by 7 | Viewed by 3092
Abstract
Digitalization is a major change driver in manufacturing and is nowadays typically linked to terms like Industry 4 [...] Full article
(This article belongs to the Special Issue Cyber Physical Production Systems)
13 pages, 17820 KiB  
Communication
Synthesis of Bulk Zr48Cu36Al8Ag8 Metallic Glass by Hot Pressing of Amorphous Powders
by Tianbing He, Nevaf Ciftci, Volker Uhlenwinkel and Sergio Scudino
J. Manuf. Mater. Process. 2021, 5(1), 23; https://doi.org/10.3390/jmmp5010023 - 9 Mar 2021
Cited by 7 | Viewed by 2832
Abstract
The critical cooling rate necessary for glass formation via melt solidification poses inherent constraints on sample size using conventional casting techniques. This drawback can be overcome by pressure-assisted sintering of metallic glass powders at temperatures above the glass transition, where the material shows [...] Read more.
The critical cooling rate necessary for glass formation via melt solidification poses inherent constraints on sample size using conventional casting techniques. This drawback can be overcome by pressure-assisted sintering of metallic glass powders at temperatures above the glass transition, where the material shows viscous-flow behavior. Partial crystallization during sintering usually exacerbates the inherent brittleness of metallic glasses and thus needs to be avoided. In order to achieve high density of the bulk specimens while avoiding (or minimizing) crystallization, the optimal combination between low viscosity and long incubation time for crystallization must be identified. Here, by carefully selecting the time–temperature window for powder consolidation, we synthesized highly dense Zr48Cu36Ag8Al8 bulk metallic glass (BMG) with mechanical properties comparable with its cast counterpart. The larger ZrCu-based BMG specimens fabricated in this work could then be post-processed by flash-annealing, offering the possibility to fabricate monolithic metallic glasses and glass–matrix composites with enhanced room-temperature plastic deformation. Full article
(This article belongs to the Special Issue Powder Metallurgy and Additive Manufacturing/3D Printing of Materials)
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Figure 1
<p>(<b>a</b>) Schematic continuous cooling transformation (CCT) diagram for a glass-forming system showing the critical cooling rate, <span class="html-italic">R</span><sub>c</sub>, for glass formation; (<b>b</b>) schematic continuous heating transformation (CHT) diagram for conventional heating rates showing the principles of consolidation of amorphous powders. <span class="html-italic">T<sub>L</sub></span> = liquidus temperature; SCL = supercooled liquid; <span class="html-italic">τ</span> = incubation time.</p>
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<p>(<b>a</b>) Particle morphology and (<b>b</b>) particle size distribution of the gas-atomized Zr<sub>48</sub>Cu<sub>36</sub>Al<sub>8</sub>Ag<sub>8</sub> glassy powder.</p>
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<p>(<b>a</b>) The hot-pressing setup and (<b>b</b>) close view of the controlled atmosphere chamber. (<b>c</b>) Schematic of the hot-pressing process.</p>
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<p>(<b>a</b>) Isochronal differential scanning calorimetry (DSC) scan (40 K/min) and viscosity vs. temperature (40 K/min) for the Zr<sub>48</sub>Cu<sub>36</sub>Al<sub>8</sub>Ag<sub>8</sub> glassy powder; inset: change of the sample height as a function of temperature; (<b>b</b>) isothermal DSC scans of the powder carried out at 673, 743, 753, and 763 K; (<b>c</b>) isochronal DSC scans (40 K/min) for the Zr<sub>48</sub>Cu<sub>36</sub>Al<sub>8</sub>Ag<sub>8</sub> glassy powder and bulk specimens hot pressed (HP) at different temperatures; (<b>d</b>) XRD patterns of as-atomized powder and hot-pressed samples.</p>
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<p>Microstructure of the Zr<sub>48</sub>Cu<sub>36</sub>Ag<sub>8</sub>Al<sub>8</sub> BMGs hot pressed at (<b>a</b>) 743 K, (<b>b</b>) 753 K, and (<b>c</b>) 758 K; (<b>d</b>) µ-CT (computed tomography) reconstruction of the Zr<sub>48</sub>Cu<sub>36</sub>Al<sub>8</sub>Ag<sub>8</sub> powder hot pressed at 753 K showing minor impurities but no porosity.</p>
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<p>(<b>a</b>) Representative room-temperature stress–strain curves under compression for the Zr<sub>48</sub>Cu<sub>36</sub>Al<sub>8</sub>Ag<sub>8</sub> bulk specimen hot pressed at 753 K and the corresponding material synthesized by casting [<a href="#B20-jmmp-05-00023" class="html-bibr">20</a>]. For clarity, the curve of HP 753 K was offset along the horizontal axis. (<b>b</b>) The five compressive stress–strain curves of the Zr<sub>48</sub>Cu<sub>36</sub>Al<sub>8</sub>Ag<sub>8</sub> BMG hot-pressed at 753 K used to derive average strength and corresponding error.</p>
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<p>(<b>a</b>) Shrinking curves as a function of the holding time for the isothermal stage of hot pressing (represented by the shadowed area) of the Zr<sub>48</sub>Cu<sub>36</sub>Al<sub>8</sub>Ag<sub>8</sub> powder at 743 and 753 K along with the crystallized fractions derived from the areas of the corresponding isothermal DSC peaks in <a href="#jmmp-05-00023-f004" class="html-fig">Figure 4</a>b. Note that the curves are shifted so that the isothermal stage starts at 0 min for both experiments. The vertical dashed line represents the time when the pressure was released. (<b>b</b>) Schematic representing the variation of contact area for a particle in an ensemble of particles with cubic packed structure and corresponding to the shrinking curves in (<b>a</b>). The sphere is progressively deformed, increasing the contact area (white circles). (<b>c</b>) Schematic representation of glass crystallization accompanying the sintering process, corresponding to the blue curves in (<b>a</b>). Crystals embedded in the glassy matrix are represented as dark blue pentagons.</p>
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<p>Surface morphology of a Zr<sub>48</sub>Cu<sub>36</sub>Al<sub>8</sub>Ag<sub>8</sub> bulk specimen consolidated at 743 K fractured upon release of the pressure during hot pressing showing the shape changes from spherical to polygonal.</p>
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13 pages, 1793 KiB  
Article
Optimization of Abrasive Flow Nano-Finishing Processes by Adopting Artificial Viral Intelligence
by Nikolaos A. Fountas and Nikolaos M. Vaxevanidis
J. Manuf. Mater. Process. 2021, 5(1), 22; https://doi.org/10.3390/jmmp5010022 - 8 Mar 2021
Cited by 12 | Viewed by 2978
Abstract
This work deals with the optimization of crucial process parameters related to the abrasive flow machining applications at micro/nano-levels. The optimal combination of abrasive flow machining parameters for nano-finishing has been determined by applying a modified virus-evolutionary genetic algorithm. This algorithm implements two [...] Read more.
This work deals with the optimization of crucial process parameters related to the abrasive flow machining applications at micro/nano-levels. The optimal combination of abrasive flow machining parameters for nano-finishing has been determined by applying a modified virus-evolutionary genetic algorithm. This algorithm implements two populations: One comprising the hosts and one comprising the viruses. Viruses act as information carriers and thus they contribute to the algorithm by boosting efficient schemata in binary coding to facilitate both the arrival at global optimal solutions and rapid convergence speed. Three cases related to abrasive flow machining have been selected from the literature to implement the algorithm, and the results corresponding to them have been compared to those available by the selected contributions. It has been verified that the results obtained by the virus-evolutionary genetic algorithm are not only practically viable, but far more promising compared to others as well. The three cases selected are the traditional “abrasive flow finishing,” the “rotating workpiece” abrasive flow finishing, and the “rotational-magnetorheological” abrasive flow finishing. Full article
(This article belongs to the Special Issue Advances in Micro and Nanomanufacturing)
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Figure 1
<p>(<b>a</b>) Transduction operation for the creation of a virus individual, (<b>b</b>) reverse transcription operation for infecting an individual with a virus, (<b>c</b>) infected individual after the reverse transcription operation performed by the virus, and (<b>d</b>) partial transduction operation for changing the virus scheme.</p>
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<p>The procedure of viral infection in the virus−evolutionary genetic algorithm.</p>
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<p>Convergence graphs for optimizing the conventional abrasive flow nano-finishing process [<a href="#B2-jmmp-05-00022" class="html-bibr">2</a>] using VEGA, under the constraint of: (<b>a</b>) <span class="html-italic">Ra_</span>max ≤ 0.7 μm, (<b>b</b>) <span class="html-italic">Ra_</span>max ≤ 0.6 μm, (<b>c</b>) <span class="html-italic">Ra_</span>max ≤ 0.5 μm, and (<b>d</b>) <span class="html-italic">Ra_</span>max ≤ 0.4 μm.</p>
Full article ">Figure 3 Cont.
<p>Convergence graphs for optimizing the conventional abrasive flow nano-finishing process [<a href="#B2-jmmp-05-00022" class="html-bibr">2</a>] using VEGA, under the constraint of: (<b>a</b>) <span class="html-italic">Ra_</span>max ≤ 0.7 μm, (<b>b</b>) <span class="html-italic">Ra_</span>max ≤ 0.6 μm, (<b>c</b>) <span class="html-italic">Ra_</span>max ≤ 0.5 μm, and (<b>d</b>) <span class="html-italic">Ra_</span>max ≤ 0.4 μm.</p>
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<p>Comparative results between the optimization objectives of min<span class="html-italic">Ra</span> and max<span class="html-italic">MRR</span> corresponding to GA, VEGA, and actual experimental results of conventional abrasive flow nano-finishing process [<a href="#B2-jmmp-05-00022" class="html-bibr">2</a>]: (<b>a</b>) min<span class="html-italic">Ra</span>, (<b>b</b>) max<span class="html-italic">MRR</span>.</p>
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<p>Convergence graphs for optimizing the rotating workpiece abrasive flow nano-finishing process [<a href="#B3-jmmp-05-00022" class="html-bibr">3</a>] using VEGA for the cases of: (<b>a</b>) Aluminum alloy, (<b>b</b>) aluminum alloy/SiC10%, and (<b>c</b>) aluminum alloy/SiC15%.</p>
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<p>Nondominated Pareto optimal solutions obtained by VEGA for optimizing the rotational-magnetorheological abrasive flow nano-finishing process [<a href="#B4-jmmp-05-00022" class="html-bibr">4</a>].</p>
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<p>Contour plots for examining the effects of independent variables for controlling the rotational-magnetorheological abrasive flow nano-finishing process, with reference to experimental results presented in [<a href="#B4-jmmp-05-00022" class="html-bibr">4</a>], (<b>a</b>) for max%Δ<span class="html-italic">Ra</span> vs. P (bar), N (cyles); (<b>b</b>) for max%Δ<span class="html-italic">Ra</span> vs. S (rpm), R (ratio).</p>
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16 pages, 39129 KiB  
Article
Turbine Blade Tip Repair by Laser Directed Energy Deposition Additive Manufacturing Using a Rene 142–MERL 72 Powder Blend
by Mohsen K. Keshavarz, Alexandre Gontcharov, Paul Lowden, Anthony Chan, Devesh Kulkarni and Mathieu Brochu
J. Manuf. Mater. Process. 2021, 5(1), 21; https://doi.org/10.3390/jmmp5010021 - 1 Mar 2021
Cited by 19 | Viewed by 5930
Abstract
Laser directed energy deposition (LDED) was used with a powder blend comprising 75 wt.% Rene 142 and 25 wt.% of Merl 72 (4275M72) for turbine blade tip repair applications. Sound samples could be deposited at ambient temperature on Haynes 230. The microstructural analyses [...] Read more.
Laser directed energy deposition (LDED) was used with a powder blend comprising 75 wt.% Rene 142 and 25 wt.% of Merl 72 (4275M72) for turbine blade tip repair applications. Sound samples could be deposited at ambient temperature on Haynes 230. The microstructural analyses showed the presence of fine gamma prime precipitates in the as-deposited samples, while after aging, the alloy possessed around 40 vol.% with a bimodal precipitate size distribution. Also, the alloy contained Ta-Hf-W carbides in different sizes and shapes. Tensile testing from room temperature up to 1366 K was performed. The 4275M72 deposits possessed higher tensile properties compared to Rene 80 in this temperature range but lower elongations at the elevated temperatures. The creep properties of 4275M72 samples at 1255 K were superior to Rene 80. Also, the oxidation resistance of deposited 4275M72 was similar to Rene 142. The combination of high mechanical properties, creep behavior, and oxidation resistance of LDEDed 4275M72 makes it a suitable alloy for tip repair of turbine blades. Full article
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Figure 1
<p>(<b>a</b>) Schematic of a typical turbine blade tip with a squealer rim, inset shows cross-section of the tip and hypothetically repaired squealer tip (<b>b</b>) schematic of LDED head with coaxial powder feeding, and (<b>c</b>) schematic of a bi-directionally deposited sample and orientation of extracted tensile specimen in respect to the build direction.</p>
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<p>PSD analysis and powder morphology of (<b>a</b>,<b>b</b>) R142, and (<b>c</b>,<b>d</b>) M72 powders.</p>
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<p>PSD analysis and powder morphology of (<b>a</b>,<b>b</b>) R142, and (<b>c</b>,<b>d</b>) M72 powders.</p>
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<p>(<b>a</b>) A multi-layer 4275M72 deposition on an H230 substrate using LDED, and (<b>b</b>) polished cross-section of deposited layers showing pores.</p>
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<p>Optical micrographs of (<b>a</b>) bi-directionally deposited sample, and (<b>b</b>) higher magnification of the selected area in panel (<b>a</b>).</p>
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<p>XRD patterns of LDEDed 4275M72 after heat treatment. The PDF numbers of matching phases are shown in parentheses.</p>
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<p>(<b>a</b>) Typical microstructural features of LDEDed 4275M72 in the as-built condition, and evolution of γ’ precipitates in (<b>b</b>–<b>d</b>) 0.5 mm from the top, (<b>e</b>–<b>g</b>) 15 mm from the top of the build (middle), and (<b>h</b>–<b>j</b>) 2 mm from the substrate interface (bottom).</p>
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<p>(<b>a</b>) A typical back-scattered electron image of aged 4275M72 deposition, and evolution of γ’ precipitates after aging in (<b>b</b>,<b>c</b>) 1 mm from the top, (<b>d</b>,<b>e</b>) 15 mm from the top of the build (middle), and (<b>f</b>,<b>g</b>) 2 mm from the substrate interface (bottom).</p>
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<p>(<b>a</b>) Definition of shape parameter ratio (η) to quantitatively evaluate γ’ shapes, and (<b>b</b>) average values of η in three sections of aged 4275M72 LDEDed sample. The error bars represent standard deviations of the values.</p>
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<p>(<b>a</b>) The 4275M72 deposits produced from the powder blend after PFHT depicting precipitation of γ’ phase and Ta-Hf-W-based carbides in the γ-matrix, (<b>b</b>–<b>d</b>) EDS spectra of phases indicated in panel (<b>a</b>), and (<b>e</b>) EDS elemental maps of carbides in panel (<b>a</b>).</p>
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<p>Vickers microhardness measurement results in as-deposited and aged 4275M72 samples along the build direction from the top layer towards the substrate.</p>
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<p>Fractographs of tensile testing samples of LDEDed 4275M72 depicting (<b>a</b>) a secondary electron (SE) image and (<b>b</b>) a backscattered (BSE) image of samples tested at room temperature, and (<b>c</b>) SE image and (<b>d</b>) BSE image of the sample tested at 1366.5 K.</p>
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<p>Larson–Miller plot of 4275M72 compared to R80 [<a href="#B37-jmmp-05-00021" class="html-bibr">37</a>].</p>
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<p>Cyclic oxidation resistance of LDEDed 4275M72 compared to R142 and R80 welds at 1120 °C (1393 K).</p>
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12 pages, 17684 KiB  
Article
Edge Grinding Characteristics of Display Glass Substrate
by Dennis Wee Keong Neo, Kui Liu, Rui Huang and Hu Wu
J. Manuf. Mater. Process. 2021, 5(1), 20; https://doi.org/10.3390/jmmp5010020 - 1 Mar 2021
Cited by 1 | Viewed by 3595
Abstract
Display glass substrate as a brittle material is very challenging to machine due to its excellent physical, mechanical, electrical, and optical properties such as high hardness, high strength, high wear resistance, good fracture toughness, good chemical stability, and good thermal stability. On the [...] Read more.
Display glass substrate as a brittle material is very challenging to machine due to its excellent physical, mechanical, electrical, and optical properties such as high hardness, high strength, high wear resistance, good fracture toughness, good chemical stability, and good thermal stability. On the basis of Griffith fracture mechanics, our theoretical analysis indicated that edge grinding of the display glass substrate is under brittle mode when grinding with the given conditions, which was verified by the experimental studies of ground glass edge surface topography and fractured surface obtained. Grinding force (Fy) in the vertical direction was much larger than grinding force (Fx) in the horizontal direction, causing a large compressive stress acting on the grinding glass edge. Grinding torque was slightly increased with the increase of grinding speed. Grinding temperature was very high when measured under dry grinding compared with measurement under high-pressure coolant. Grinding of glass substrate edge was performed partially under ductile mode machining in the experimental conditions, which can be attributed to and contributed by those micro cutting edges generated by the fractured diamond grit on the grinding wheel surface. Full article
(This article belongs to the Special Issue Progress in Precision Machining)
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<p>Close view of the edge grinding experimental setup.</p>
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<p>Data acquisition system for edge grinding of display glass substrates.</p>
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<p>Experimental details of temperature measurement in edge grinding of glass substrates: (<b>a</b>) schematic illustration of embedded thermocouples; (<b>b</b>) actual embedded thermocouples in glass substrates.</p>
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<p>Grinding forces and spindle torque measured in the plane grinding of glass substrate edges: (<b>a</b>) grinding force, <span class="html-italic">F<sub>x</sub></span>; (<b>b</b>) grinding force, <span class="html-italic">F<sub>y</sub></span>; (<b>c</b>) grinding spindle torque.</p>
Full article ">Figure 4 Cont.
<p>Grinding forces and spindle torque measured in the plane grinding of glass substrate edges: (<b>a</b>) grinding force, <span class="html-italic">F<sub>x</sub></span>; (<b>b</b>) grinding force, <span class="html-italic">F<sub>y</sub></span>; (<b>c</b>) grinding spindle torque.</p>
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<p>Effects of grinding distance and speed on grinding force and torque: (<b>a</b>) grinding force vs. grinding distance; (<b>b</b>) grinding force vs. spindle speed; (<b>c</b>) grinding torque vs. grinding distance.</p>
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<p>Grinding temperature measured in plane grinding of glass substrate edges: (<b>a</b>) grinding under high-pressure coolant; (<b>b</b>) grinding under dry conditions.</p>
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<p>Ground edge surface topography and defects: (<b>a</b>) overview of ground edge; (<b>b</b>) close view of ground edge; (<b>c</b>) chips adhered on the ground edge; (<b>d</b>) defects on ground edge.</p>
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<p>Diamond grinding wheel surface topography before and after usage: (<b>a</b>) fresh grinding wheel; (<b>b</b>) used grinding wheel.</p>
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<p>Diamond grinding wheel wear patterns: (<b>a</b>) a protruded diamond grit; (<b>b</b>) a worn diamond grit; (<b>c</b>) diamond grit pull out.</p>
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<p>Ground surfaces of glass substrate edges: (<b>a</b>) fractured surface obtained under the spindle speed of 3000 rpm, feed rate of 7000 mm/min, and depth of cut of 0.1 mm; (<b>b</b>) smooth surface obtained under the spindle speed of 6000 rpm, feed rate of 5 mm/min, and depth of cut of 0.05 mm.</p>
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14 pages, 9922 KiB  
Article
Manufacturing Concept and Prototype for Train Component Using the FSW Process
by Elizabeth Hoyos, Santiago Escobar, Jeroen De Backer, Jonathan Martin and Mauricio Palacio
J. Manuf. Mater. Process. 2021, 5(1), 19; https://doi.org/10.3390/jmmp5010019 - 13 Feb 2021
Cited by 5 | Viewed by 2744
Abstract
Friction stir welding (FSW) is a process originally developed for joining light materials, such as aluminum and magnesium, as an answer to their poor weldability by conventional fusion processes. In Colombia, the technique has been studied but its industrial implementation is uncommon, due [...] Read more.
Friction stir welding (FSW) is a process originally developed for joining light materials, such as aluminum and magnesium, as an answer to their poor weldability by conventional fusion processes. In Colombia, the technique has been studied but its industrial implementation is uncommon, due to the high cost of specialized machinery and the unfamiliarity with the technique of local industries. This article presents an implementation case study of FSW on a 6082-aluminum alloy train component from Metro de Medellín (MdM), aiming to establish the component design changes required to accommodate the FSW process, and conventional machines available in the local area which may be available for welding. Additionally, a simple comparison was made between the cost of this approach versus the manufacturing strategy currently used for the selected component. Initially, welding forces were measured when performing the seam on the selected component using an FSW machine. This data was then used to downselect the local milling machines with these capabilities. A simple but specific tool was designed for the geometry of one of the component features. Finally, a prototype was fabricated, and weld samples were obtained, polished, etched, and examined using a microhardness machine and an optical microscope. Results show a good opportunity for the execution of simple components with uniform geometries, which can be carried out using locally available machinery because they do not surpass their maximum loading capacity, the welds do not present visible discontinuities, and an average hardness of 69.5 HV and mechanical efficiency of 95% can be achieved. Additionally, the manufacturing process is around 30% cheaper compared to traditional methods, making the application viable, economically speaking. Full article
(This article belongs to the Special Issue Friction Stir Welding and Related Technologies)
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<p>Component used for the study case. (<b>a</b>) Extrusion, (<b>b</b>) bracket with welding illustration, (<b>c</b>) lid with welding illustration and (<b>d</b>) location in the wagon.</p>
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<p>Geometry of the tool used for welding. (<b>a</b>) General perspective, (<b>b</b>) detail on shoulder and pin.</p>
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<p>Bracket insert. (<b>a</b>) Machined piece and (<b>b</b>) weld path (red).</p>
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<p>Lid proposed redesign using bolted assembly.</p>
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<p>C-shaped channel configuration. (<b>a</b>) Transverse section of the two L-shaped profiles, (<b>b</b>) full schematic, (<b>c</b>) weld configuration downwards and (<b>d</b>) weld configuration upwards.</p>
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<p>Different welds performed. (<b>a</b>) Longitudinal weld, (<b>b</b>) bracket weld.</p>
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<p>Preliminary “C” joints. (<b>a</b>) Downwards configuration and (<b>b</b>) upwards configuration.</p>
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<p>Axial force against weld length.</p>
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<p>(<b>a</b>) Different paths employed, (<b>b</b>) actual weld on the C-shaped profile.</p>
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<p>(<b>a</b>) Longitudinal weld performed in the C shape, (<b>b</b>) advancing side and (<b>c</b>) retreating side of the bracket weld.</p>
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<p>(<b>a</b>) Top and (<b>b</b>) bottom close-up of the bracket weld in the advancing side.</p>
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<p>Hardness measurement positions.</p>
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<p>Horizontal hardness measurements.</p>
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<p>Final prototype images, (<b>a</b>) top view, (<b>b</b>) side view, (<b>c</b>) close-up of the bracket feature and (<b>d</b>) close-up picture of the bracket’s weld.</p>
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14 pages, 7072 KiB  
Article
Numerical and Experimental Investigation of the Impact of the Electromagnetic Properties of the Die Materials in Electromagnetic Forming of Thin Sheet Metal
by Björn Beckschwarte, Lasse Langstädtler, Christian Schenck, Marius Herrmann and Bernd Kuhfuss
J. Manuf. Mater. Process. 2021, 5(1), 18; https://doi.org/10.3390/jmmp5010018 - 12 Feb 2021
Cited by 10 | Viewed by 2406
Abstract
In electromagnetic forming of thin sheet metal, the die is located within the effective range of the electromagnetic wave. Correspondingly, a current is induced not only in the sheet metal, but also in the die. Like the current in the workpiece, also the [...] Read more.
In electromagnetic forming of thin sheet metal, the die is located within the effective range of the electromagnetic wave. Correspondingly, a current is induced not only in the sheet metal, but also in the die. Like the current in the workpiece, also the current in the die interacts with the electromagnetic wave, resulting in Lorentz forces and changes of the electromagnetic field. With the aim to study the influence of different electromagnetic die properties in terms of specific electric resistance and relative magnetic permeability, electromagnetic simulations were carried out. A change in the resulting forming forces in the sheet metals was determined. To confirm the simulation results, electromagnetic forming and embossing tests were carried out with the corresponding die materials. The results from simulation and experiment were in good agreement. Full article
(This article belongs to the Special Issue Impulse-Based Manufacturing Technologies)
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<p>Simulation model; (<b>a</b>) schematic; (<b>b</b>) input current (tool coil).</p>
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<p>Experimental setup; (<b>a</b>) equivalent circuit diagram; (<b>b</b>) tool coil.</p>
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<p>Forming experiment; (<b>a</b>) free forming die geometry; (<b>b</b>) resulting workpiece geometry.</p>
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<p>Embossing experiment; (<b>a</b>) die geometry; (<b>b</b>) resulting workpiece geometries.</p>
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<p>Simulated impulse as a function of the resulting skin depth for the parameter combination of specific resistance <span class="html-italic">ρ</span> and relative permeability <span class="html-italic">μ<sub>r</sub></span> values.</p>
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<p>Simulated impulse <span class="html-italic">J</span> on thin aluminium sheets for different electromagnetic properties of the die material.</p>
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<p>Maximum current density <span class="html-italic">j<sub>max</sub></span> in the workpiece volume in respect of the specific resistance <span class="html-italic">ρ</span> of the die material.</p>
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<p>Integral of the current density j in the die volume over time in respect of the specific resistance <span class="html-italic">ρ</span> of the die material.</p>
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<p>Integral of the negative force components over time in respect of the specific resistance <span class="html-italic">ρ</span> of the die material.</p>
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<p>Distribution of the magnetic field with high relative permeability <span class="html-italic">μ<sub>r</sub></span> of the die during the maximum current density in the tool coil.</p>
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<p>Distribution of the magnetic field with low relative permeability <span class="html-italic">μ<sub>r</sub></span> of the die during the maximum current density in the tool coil.</p>
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<p>Average magnetic field strength H<sub>a</sub> in the workpiece in respect of the specific resistance <span class="html-italic">ρ</span> of the die material.</p>
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<p>Field strength deviation H<sub>d</sub> (standard deviation of the magnetic field strength) in the workpiece in respect of the specific resistance of the die material.</p>
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<p>Corresponding experimental bulge height h for the simulated impulse <span class="html-italic">J</span> for different die materials with linear correlation line for a 1250 J forming experiment.</p>
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<p>Comparison of the inner corner radius in two directions and the embossing result of two embossing dies made of austenitic chrome-nickel steel and copper, cf. [<a href="#B25-jmmp-05-00018" class="html-bibr">25</a>].</p>
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15 pages, 8462 KiB  
Article
Fundamental Investigation of Diamond Cutting of Micro V-Shaped Grooves on a Polycrystalline Soft-Brittle Material
by Weihai Huang and Jiwang Yan
J. Manuf. Mater. Process. 2021, 5(1), 17; https://doi.org/10.3390/jmmp5010017 - 8 Feb 2021
Cited by 11 | Viewed by 3583
Abstract
Fabricating micro-structures on optical materials has received great interest in recent years. In this work, micro-grooving experiments were performed on polycrystalline zinc selenide (ZnSe) to investigate the feasibility of surface micro-structuring on polycrystalline soft-brittle material by diamond turning. A photosensitive resin was coated [...] Read more.
Fabricating micro-structures on optical materials has received great interest in recent years. In this work, micro-grooving experiments were performed on polycrystalline zinc selenide (ZnSe) to investigate the feasibility of surface micro-structuring on polycrystalline soft-brittle material by diamond turning. A photosensitive resin was coated on the workpiece before cutting, and it was found that the coating was effective in suppressing brittle fractures at the edges of the grooves. The effect of tool feed rate in groove depth direction was examined. Results showed that the defect morphology on the groove surface was affected by the tool feed rate. The crystallographic orientation of grains around the groove was characterized by electron backscatter diffraction (EBSD), and it was found that the formation of defects was strongly dependent on the angle of groove surface with respect to the cleavage plane of grain. The stress distribution of the micro-grooving process was investigated by the finite element method. Results showed that the location of tensile stresses in the coated workpiece was farther from the edge of the groove compared with that in the uncoated workpiece, verifying the experimental result that brittle fractures were suppressed by the resin coating. Full article
(This article belongs to the Special Issue Progress in Precision Machining)
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<p>Machining models of (<b>a</b>) flat surface cutting using a round-nosed tool; (<b>b</b>) V-shaped groove cutting using a V-shaped tool.</p>
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<p>(<b>a</b>) Experimental setup for off-axis diamond cutting of micro V-shaped grooves; (<b>b</b>) SEM image of the diamond tool with an extremely sharpened tool tip.</p>
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<p>Schematic diagram of the experimental procedure: (<b>a</b>) Off-axis cutting; (<b>b</b>) cutting grooves on uncoated workpiece; (<b>c</b>) coating workpiece surface with photosensitive resin; (<b>d</b>) cutting grooves on coated workpiece; and (<b>e</b>) removing the photosensitive resin. The workpiece shown in <a href="#jmmp-05-00017-f003" class="html-fig">Figure 3</a><b>b</b>–<b>d</b> is a small-area extraction from <a href="#jmmp-05-00017-f003" class="html-fig">Figure 3</a>a. For simplicity, the grooves in <a href="#jmmp-05-00017-f003" class="html-fig">Figure 3</a><b>b</b>–<b>d</b> are approximated to be straight, as the curvature is extremely small in such a small area.</p>
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<p>SEM images of the V-shaped groove before removing the resin coating (<b>a</b>) general view; (<b>b</b>) close-up view. (Cutting parameters: <span class="html-italic">f<sub>z</sub></span> = 10 nm/rev, and <span class="html-italic">d</span> = 9 μm).</p>
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<p>SEM images of the V-shaped grooves after removing the resin coating (all grooves were cut at <span class="html-italic">f<sub>z</sub></span> = 10 nm/rev).</p>
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<p>SEM images of the V-shaped grooves after removing the resin coating (all grooves were cut at <span class="html-italic">f<sub>z</sub></span> = 10 nm/rev).</p>
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<p>SEM images of V-shaped grooves cut at different feed rates: (<b>a</b>) 20 nm/rev; (<b>b</b>) 40 nm/rev; (<b>c</b>) 60 nm/rev.</p>
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<p>Schematic diagram of double-groove cutting experiment.</p>
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<p>(<b>a</b>) SEM image of a typical area in the double groove taken in Inlens Duo mode; (<b>b</b>) schematic drawings of the grain boundary plane which corresponds to the grain boundary in SEM image; (<b>c</b>) IPF map of the region outlined by red box in SEM image; (<b>d</b>) misorientation angle profile along Line <span class="html-italic">PP</span>’ indicated in IPF map. Schematic diagram of the spatial relationship between V-shaped groove and (<b>e</b>) grain A and (<b>f</b>) grain B.</p>
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<p>SEM image of the same region as <a href="#jmmp-05-00017-f008" class="html-fig">Figure 8</a>a by HE-SE detector.</p>
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<p>Schematic model for (<b>a</b>) formation of submicron-pits; (<b>b</b>) formation of micron-crater.</p>
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<p>(<b>a</b>) 3D simulation model of microgrooving process; (<b>b</b>) schematic of simulation configuration showing two paths of cutting tool.</p>
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<p>Schematic diagram of crack generation in cutting groove on brittle materials.</p>
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<p>FEM-simulated tensile stress along cutting direction (<span class="html-italic">y</span>-axis direction) in: (<b>a</b>) Uncoated workpiece; (<b>b</b>) coated workpiece.</p>
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<p>FEM-simulated tensile stress along the direction normal to the workpiece surface (<span class="html-italic">z</span>-axis direction) in: (<b>a</b>) Uncoated workpiece; (<b>b</b>) coated workpiece.</p>
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25 pages, 41821 KiB  
Article
Shielded Active Gas Forge Welding of an L80 Steel in a Small Scale Shielded Active Gas Forge Welding Machine
by Vinothkumar Palanisamy, Jan Ketil Solberg and Per Thomas Moe
J. Manuf. Mater. Process. 2021, 5(1), 16; https://doi.org/10.3390/jmmp5010016 - 8 Feb 2021
Cited by 3 | Viewed by 3235 | Correction
Abstract
The Shielded Active Gas Forge Welding (SAG-FW) method is a solid-state welding technique in which the mating surfaces are heated by induction heating or direct electrical heating before being forged together to form a weld. In this article, an API 5CT L80 grade [...] Read more.
The Shielded Active Gas Forge Welding (SAG-FW) method is a solid-state welding technique in which the mating surfaces are heated by induction heating or direct electrical heating before being forged together to form a weld. In this article, an API 5CT L80 grade carbon steel alloy has been welded using the SAG-FW method. A small-scale forge welding machine has been used to join miniature pipes extracted from a large pipe wall. The welding was performed at three different forging temperatures, i.e., 1300 °C, 1150 °C and 950 °C, in some cases followed by one or two post weld heat treatment cycles. In order to qualify the welds, mechanical and corrosion testing was performed on miniature samples extracted from the welded pipes. In addition, the microstructure of the welds was analysed, and electron probe microanalysis was carried out to control that no oxide film had formed along the weld line. Based on the complete set of experimental results, promising parameters for SAG-FW welding of the API 5CT L80 grade steel are suggested. The most promising procedure includes forging at relative high temperature (1150 °C) followed by rapid cooling and a short temper. This procedure was found to give a weld zone microstructure dominated by tempered martensite with promising mechanical and corrosion properties. The investigation confirmed that small scale forge welding testing is a useful tool in the development of welding parameters for full size SAG-FW welding. Full article
(This article belongs to the Special Issue Metal Forming and Joining)
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<p>(<b>a</b>) Mating end dimensions of miniature pipes, (<b>b</b>) schematic drawing of welding set-up and process.</p>
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<p>Miniature Charpy V-notch specimen. Weld interface is located vertically below the notch.</p>
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<p>Length section of a welded pipe etched with 2% Nital, WZ = Weld Zone, TZ = Transition Zone, ICZ = InterCritical Zone (base metal heated to a temperature between Ac<sub>1</sub> and Ac<sub>3</sub>), BM = Base Metal.</p>
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<p>Temperature gradient along the longitudinal axis of the pipes prior to forging for different welding temperatures.</p>
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<p>WDS oxygen K<sub>α</sub> line scan across the weld line of specimen W9.</p>
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<p>Weld zone microstructures of pipes forge welded at 1300 °C. The cooling rates below 900 °C and additional heat treatment cycles are given at the bottom of the micrographs. The weld line is located vertically in the middle of each micrograph, in most cases not visible. Martensite is coloured light brown.</p>
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<p>Weld zone microstructures of pipes forge welded at 1150 °C (<b>a</b>–<b>d</b>) and 950 °C (<b>e</b>). The cooling rates below 900 °C and additional heat treatment cycles are given at the bottom of the micrographs. The weld line is located vertically in the middle of each micrograph, in most cases not visible. Martensite is coloured light brown.</p>
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<p>HAZ microstructures recorded 5 mm away from the weld line. (<b>a</b>) Very fine-grained ferrite in weld W10, (<b>b</b>) ferrite/pearlite in weld W5, (<b>c</b>) Mixture of relatively coarse and very fine ferrite in weld W3, probably ICZ microstructure.</p>
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<p>Hardness profiles of pipes forge welded at 1300 °C followed by different thermal conditions. Base metal hardness is reached at a distance of 7.5 mm from weld line.</p>
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<p>Hardness profiles of pipes forge welded at 1150 °C (W2, W8, W10, W11) and 950 °C (W3) followed by different thermal conditions. Base metal hardness is reached at a distance of 6–7.5 mm from the weld line.</p>
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<p>Insufficient material flow in the outer weld cap of the pipe that was welded at 950 °C (W3), resulting in a camel shaped outer cap.</p>
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<p>Typical face bent specimen.</p>
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<p>Fractured bend test specimen, forged welded at 1300 °C followed by cooling at10 °C/s (W4).</p>
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<p>Comparison of 0 °C Charpy V values of full size (3 parallels) and miniature size (2 parallels) SMITWELD thermally simulated specimens. The peak temperature during the simulations was 1300 °C, and cooling was done as described in <a href="#sec2-jmmp-05-00016" class="html-sec">Section 2</a>. Each graph includes two specimens that had been tempered for 1 s at 680 °C after having been cooled at respectively 10 °C/s and 60 °C/s from 920 °C. Base metal values are also given. The full size values are taken from Ref. [<a href="#B14-jmmp-05-00016" class="html-bibr">14</a>].</p>
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<p>IFM scanned image (<b>top</b>) and topography analysis (<b>bottom</b>) of corroded specimen W1.</p>
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<p>IFM topography analysis of corroded specimen W8. (There is no correlation between the depth values in <a href="#jmmp-05-00016-f015" class="html-fig">Figure 15</a> and <a href="#jmmp-05-00016-f016" class="html-fig">Figure 16</a> since there was no common reference surface during the recordings).</p>
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18 pages, 7024 KiB  
Article
Die Material Selection Criteria for Aluminum Hot Stamping
by Maider Muro, Ines Aseguinolaza and Garikoitz Artola
J. Manuf. Mater. Process. 2021, 5(1), 15; https://doi.org/10.3390/jmmp5010015 - 2 Feb 2021
Cited by 6 | Viewed by 4023
Abstract
The aim of this work is to develop a die material selection criterion for aluminum hot stamping applications. The criterion has been based on the back-to-back comparison of a set of reciprocating friction and wear tests. Three representatives belonging to different stamping die [...] Read more.
The aim of this work is to develop a die material selection criterion for aluminum hot stamping applications. The criterion has been based on the back-to-back comparison of a set of reciprocating friction and wear tests. Three representatives belonging to different stamping die material families have been selected for the study: a cold work steel, a hot work steel, and a cast iron. These tool materials have been combined with an exemplary member from two heat treatable aluminum families: 2XXX and 6XXX. Each die-material/aluminum–alloy combination has been tested at three temperatures: 40, 200, and 450 °C. The temperatures have been selected according to different stamping scenarios: long takt time press quenching, short takt time press quenching, and very short takt time hot forming without quenching, respectively. The results show that, among the three die material options available, the cold work steel turned out to be the most favorable option for high volume production and long takt time, the hot work steel fitted best for high volume production coupled with short takt time, and cast iron turned to outstand for short runs with prototype dies and for hot stamping without die quenching. Full article
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<p>Microstructure of the three tool material batches employed for this study, (<b>a</b>) 1.2379 steel deep etched to reveal carbide bands, (<b>b</b>) QRO90 steel etched to reveal the tempered martensite, (<b>c</b>) GJS-700-2L cast iron etched to reveal the ferrite to pearlite ratio, and (<b>d</b>) GJS-700-2L iron as-polished to reveal graphite distribution.</p>
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<p>Microstructure of the three tool material batches employed for this study, (<b>a</b>) 1.2379 steel deep etched to reveal carbide bands, (<b>b</b>) QRO90 steel etched to reveal the tempered martensite, (<b>c</b>) GJS-700-2L cast iron etched to reveal the ferrite to pearlite ratio, and (<b>d</b>) GJS-700-2L iron as-polished to reveal graphite distribution.</p>
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<p>The microstructure of the two pin materials employed in this study, (<b>a</b>) LOM micrograph of the 2011 (radial section of the bar), (<b>b</b>) LOM micrograph of the 6026 (radial section of the bar), (<b>c</b>) SEM micrograph of the 2011 with phase identification confirmed by EDX (longitudinal section of the bar), and (<b>d</b>) SEM micrograph of the 6026 with phase identification confirmed by EDX (longitudinal section of the bar).</p>
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<p>Upper and lower specimens in the SRV machine.</p>
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<p>COF evolution for each tool material–aluminium alloy combination. The graphs have been grouped per testing temperature.</p>
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<p>COF evolution for each tool material–aluminium alloy combination. The graphs have been grouped per testing temperature.</p>
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<p>Wear tracks for disks tested against 2011 pins.</p>
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<p>Wear tracks for disks tested against 6026 pins.</p>
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<p>Depth profile of the track. Tool steel vs. cast iron wear profile comparison at 40 °C.</p>
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<p>Depth profile of the track. Comparison between GJS-700-L2 disks whose COF was significantly different at 40 °C depending on the pin material.</p>
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<p>Depth profile of the track. Comparison between GJS-700-L2 disk profiles for tests against 2011 and 6026 at 200 °C.</p>
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<p>Depth profile of the track. Comparison of tool steel galling vs. cast iron wear when using 6026 pins in tests at 200 °C.</p>
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<p>Depth profile of the track. Tests carried out at 450 °C with GJS-700-2L showing how wear rules over galling both with 2011 and 6026.</p>
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<p>Depth profile of the track. Test carried out at 450 °C showing how QRO90<sup>®</sup> disks suffered galling with 6026 pins, but adhesion was much lower with 2011 pins.</p>
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<p>Depth profile of the track. Test carried out at 450 °C with 1.2379 and both 2011 and 6026.</p>
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4 pages, 178 KiB  
Editorial
Acknowledgment to Reviewers of JMMP in 2020
by JMMP Editorial Office
J. Manuf. Mater. Process. 2021, 5(1), 14; https://doi.org/10.3390/jmmp5010014 - 28 Jan 2021
Viewed by 1439
Abstract
Peer review is the driving force of journal development, and reviewers are gatekeepers who ensure that JMMP maintains its standards for the high quality of its published papers [...] Full article
14 pages, 4647 KiB  
Article
Investigation of Mechanical Loads Distribution for the Process of Generating Gear Grinding
by Patricia de Oliveira Teixeira, Jens Brimmers and Thomas Bergs
J. Manuf. Mater. Process. 2021, 5(1), 13; https://doi.org/10.3390/jmmp5010013 - 27 Jan 2021
Cited by 3 | Viewed by 2518
Abstract
In grinding, interaction between the workpiece material and rotating abrasive tool generates high thermo-mechanical loads in the contact zone. If these loads reach critically high values, workpiece material properties deteriorate. To prevent the material deterioration, several models for thermomechanical analysis of grinding processes [...] Read more.
In grinding, interaction between the workpiece material and rotating abrasive tool generates high thermo-mechanical loads in the contact zone. If these loads reach critically high values, workpiece material properties deteriorate. To prevent the material deterioration, several models for thermomechanical analysis of grinding processes have been developed. In these models, the source of heat flux is usually considered as uniform in the temperature distribution calculation. However, it is known that heat flux in grinding is generated from frictional heating as well as plastic deformation during the interaction between workpiece material and each grain from the tool. To consider these factors in a future coupled thermomechanical model specifically for the process of gear generating grinding, an investigation of the mechanical load distribution during interaction between grain and workpiece material considering the process kinematics is first required. This work aims to investigate the influence of process parameters as well as grain shape on the distribution of the mechanical loads along a single-grain in gear generating grinding. For this investigation, an adaptation of a single-grain energy model considering the chip formation mechanisms is proposed. The grinding energy as well as normal force can be determined either supported by measurements or solely based on prediction models. Full article
(This article belongs to the Special Issue Progress in Precision Machining)
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<p>Distribution of energy during the process of grinding.</p>
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<p>Chip formation mechanisms for the process of grinding.</p>
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<p>Objective and approach of the work.</p>
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<p>Description of experimental procedure and design of experiment.</p>
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<p>Modeling of friction energy.</p>
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<p>Modeling of plowing energy.</p>
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<p>Modeling of shearing energy.</p>
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<p>Validation of the adapted Linke energy model.</p>
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<p>Influence of opening angle α on the energy of each chip formation mechanism.</p>
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<p>Influence of apex angle β on the energy of each chip formation mechanism.</p>
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<p>Influence of feed rate on the energy of each chip formation mechanism.</p>
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10 pages, 4141 KiB  
Article
Micro-Injection Molding of Diffractive Structured Surfaces
by Ann-Katrin Boinski, Barnabas Adam, Arne Vogelsang, Lars Schönemann, Oltmann Riemer and Bernhard Karpuschewski
J. Manuf. Mater. Process. 2021, 5(1), 12; https://doi.org/10.3390/jmmp5010012 - 22 Jan 2021
Viewed by 2665
Abstract
In recent years, the use of highly functional optical elements has made its way into our everyday life. Its applications range from use in utility items such as cell phone cameras up to security elements on banknotes or production goods. For this purpose, [...] Read more.
In recent years, the use of highly functional optical elements has made its way into our everyday life. Its applications range from use in utility items such as cell phone cameras up to security elements on banknotes or production goods. For this purpose, the Leibniz Institute for Materials Engineering (IWT) has been developing a cutting process for the fast and cost-effective production of hologram-based diffractive optical elements. In contrast to established non-mechanical manufacturing processes, such as laser lithography or chemical etching, which are able to produce optics in large quantities and with high accuracy, the diamond turning approach is extending these properties by offering several degrees of freedom. This allows for an almost unlimited geometric complexity and a structured area of considerable size (several tenth square millimeters), achieved in a single process step. In order to introduce diffractive security features to the mass market and to actual production goods, a high-performance replication process is required as the consecutive development step. Micro injection molding represents a feasible and promising option here. In particular, diamond machining enables the integration of safety features directly into the mold insert. Not only does this make additional assembly obsolete, but the safety feature can also be placed inconspicuously in the final product. In this paper, the potential of micro-injection molding as a replication process for diffractive structured surfaces will be investigated and demonstrated. Furthermore, the optical functionality after replication will be verified and evaluated. Full article
(This article belongs to the Special Issue Progress in Precision Machining)
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<p>Process principle of diamond machining of diffractive structured surfaces. (<b>a</b>) Workpiece and nFTS with diamond tool, (<b>b</b>) structuring tool path in cutting direction, (<b>c</b>) structuring tool path in feed direction, (<b>d</b>) rectangular shaped diamond tools with a tool width of 12.04 µm.</p>
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<p>Mold design of the injection mold with mold insert.</p>
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<p>Comparison of mold (<b>a</b>,<b>c</b>) and replicated surfaces (<b>b</b>,<b>d</b>) with test structure: atomic force microscopic (AFM) measurement and cross section (<b>a</b>,<b>b</b>) and coherence scanning interferometric (CSI) measurements (<b>c</b>,<b>d</b>).</p>
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<p>Comparison of mold and replication surface structured with a test structure, (<b>a</b>,<b>b</b>) white light interferometric measurements and (<b>c</b>,<b>d</b>) cross sections. ①, ② and ③ indicate exemplary positions for cross referencing between the measurements (<b>c</b>–<b>e</b>).</p>
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<p>Atomic force microscopic measurement and cross section of the replication of a diffractive structured mold insert.</p>
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<p>(<b>a</b>) Setup for the optical testing and (<b>b</b>) reconstruction of coated replicated surface.</p>
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<p>Conformity of original and reconstructed images.</p>
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12 pages, 4086 KiB  
Article
An Efficient Ultraprecision Machining System Automating Setting Operations of Roughly Machined Workpiece
by Meng Xu, Keiichi Nakamoto and Yoshimi Takeuchi
J. Manuf. Mater. Process. 2021, 5(1), 11; https://doi.org/10.3390/jmmp5010011 - 19 Jan 2021
Cited by 5 | Viewed by 2731
Abstract
Ultraprecision machining is required in many advanced fields. To create precise parts for realizing their high performance, the whole machining process is usually conducted on the same ultraprecision machine tool to avoid setting errors by reducing setting operations. However, feed rate is relatively [...] Read more.
Ultraprecision machining is required in many advanced fields. To create precise parts for realizing their high performance, the whole machining process is usually conducted on the same ultraprecision machine tool to avoid setting errors by reducing setting operations. However, feed rate is relatively slow and machining efficiency is not so high compared to ordinary machine tools. Thus, the study aims to develop an efficient ultraprecision machining system including an industrial robot to avoid manual setting and to automate the setting operations. In this system, ultraprecision machining is conducted for the workpiece having a shape near the target shape, which is beforehand prepared by ordinary machine tools and is located on the machine table by means of an industrial robot. Since the setting errors of the roughly machined workpiece deteriorate machining accuracy, the differences from the ideal position and attitude are detected with a contact type of on-machine measurement device. Numerical control (NC) data is finally modified to compensate the identified workpiece setting errors to machine the target shape on an ultraprecision machine tool. From the experimental results, it is confirmed that the proposed system has the possibility to reduce time required in ultraprecision machining to create precise parts with high efficiency. Full article
(This article belongs to the Special Issue Progress in Precision Machining)
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<p>Experimental setup used in this study: (<b>a</b>) Five-axis controlled ultraprecision machining center ROBOnano α-0iB (FANUC corp.). (<b>b</b>) Six-axis vertical articulated industrial robot LR Mate 200iD (FANUC corp.).</p>
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<p>A non-rotational diamond tool is mounted on a B axis table via an angle plate and an on-machine measurement device is located on beside the angle plate. A probe of on-machine measurement device with R0.25 ruby ball detects the displacement while following workpieces which are set on a vacuum chuck on the C axis table.</p>
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<p>Workpiece setting errors are expressed as the disagreement between ideal and actual workpiece coordinate systems, which are established based on the ideal and actual position of workpieces, respectively. The workpiece setting errors are categorized as the position error <b><span class="html-italic">δ</span></b> or the rotation errors α, β, and γ.</p>
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<p>(<b>a</b>) Specific spheres that were created by rough cutting using a ball-end mill were arranged on the workpiece; the center point of the sphere that was adopted as the reference was calculated by detecting the spherical surface; (<b>b</b>) sphere surface was scanned in two different directions and its center position was estimated by the least squares method; (<b>c</b>) probe was fed in the −Z<sub>w</sub> direction of the ideal workpiece coordinate system until the probe touched the spherical surface, to find the bottom point of the spherical surface, and the sphere center was calculated from this bottom point and sphere radius R, which was the average radius of three different spheres.</p>
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<p>Establishment of the actual workpiece coordinate system: (<b>a</b>) Each axis of the actual workpiece coordination system was based on the three references that were the center points of specific spheres and the reference point S<sub>1</sub> was set as the origin; (<b>b</b>) rotation error α is the angle from Z<sub>w</sub>’ axis to X<sub>w</sub>-Y<sub>w</sub> plane and the position error δ is the offset of S<sub>1</sub> from its ideal position.</p>
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<p>Estimation of C axis rotational center by calculating the midpoint of the bottom points of the specific sphere before and after C axis was rotated by 180 degrees.</p>
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<p>(<b>a</b>) The target shape consists of four linear grooves having 1 mm length and crossing the same point; (<b>b</b>) grooves are machined by feeding the diamond tool along Y<sub>w</sub>’ direction where the C axis table is rotated by 90 degrees.</p>
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<p>(<b>a</b>) Microscopic images of the machined shapes before and after workpiece setting errors compensation; (<b>b</b>) Depths of the end of machined grooves measured with microscope before and after compensation.</p>
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<p>(<b>a</b>) Dimension of a diamond crown as the target shape; (<b>b</b>) dimension of the roughly machined workpiece that was 10 μm larger than the targeted size to make sure that the machining result contains the target shape safely.</p>
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<p>Tool paths for the ultraprecision machining: (<b>a</b>) Top surface was machined along Y<sub>w</sub>’ direction; (<b>b</b>) side surfaces were machined by using Y<sub>w</sub>’ and Z<sub>w</sub>’ axes at the same time to make the tool paths parallel to the side surfaces where the C axis table was rotated by 45 degrees.</p>
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<p>(<b>a</b>) Microscopic image of the machined shape before ultraprecision machining; (<b>b</b>) microscopic image of the machined shape after ultraprecision machining before workpiece setting errors compensation; (<b>c</b>) microscopic image of the machined shape after ultraprecision machining after compensation.</p>
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<p>Measured parameters of the machined shape: (<b>a</b>) The distances of two opposite edges; (<b>b</b>) the inclination of the side surface.</p>
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13 pages, 5422 KiB  
Article
Numerical and Experimental Study of AlSi Coating Effect on Nugget Size Growth in Resistance Spot Welding of Hot-Stamped Boron Steels
by Ali Afzal, Mohsen Hamedi and Chris Valentin Nielsen
J. Manuf. Mater. Process. 2021, 5(1), 10; https://doi.org/10.3390/jmmp5010010 - 15 Jan 2021
Cited by 5 | Viewed by 2669
Abstract
In recent years, increasing automotive safety by improving crashworthiness has been a focal point in the automotive industry, employing high-strength steel such as press hardenable steel (PHS). In addition to the improved strength of individual parts in the body of the vehicle, the [...] Read more.
In recent years, increasing automotive safety by improving crashworthiness has been a focal point in the automotive industry, employing high-strength steel such as press hardenable steel (PHS). In addition to the improved strength of individual parts in the body of the vehicle, the strength of the resistance-spot-welded joints of these parts is highly important to obtain a safe structure. In general, dimensions of weld nuggets are regarded as one of the criteria for the quality of spot-welded joints. In the presented research, a three-dimensional axisymmetric finite element model is developed to predict the nugget formation in resistance spot welding (RSW) of two types of PHS: the uncoated and AlSi-coated 1.8 mm boron steel after hot stamping. A fully coupled electro-thermo-mechanical analysis was conducted using the commercial software package Abaqus. The FE predicted weld nugget development is compared with experimental results. The computed weld nugget sizes show good agreement with experimental values. Full article
(This article belongs to the Special Issue Metal Forming and Joining)
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<p>Photographs for thermal processing of the samples (<b>a</b>) Putting specimens in a furnace; (<b>b</b>) Setting the temperature to 910 °C; (<b>c</b>) Taking the austenitized specimens out of the furnace after 6 min; (<b>d</b>) Specimens after quenching.</p>
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<p>Microstructure of material and appearance of AlSi coating (<b>a</b>) as-received condition; (<b>b</b>) after heat treatment similar to the hot-stamped condition.</p>
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<p>Measurement setup for electrical contact resistance.</p>
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<p>Initial mesh of the developed model.</p>
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<p>Boundary conditions in the axisymmetric resistance spot welding (RSW) model.</p>
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<p>A typical temperature distribution of the weldment at the end of RSW of uncoated steel. The scale bar shows the temperature in degrees Celsius.</p>
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<p>A typical temperature distribution of the weldment at the end of RSW of AlSi-coated steel. The scale bar shows the temperature in degrees Celsius.</p>
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<p>A representative comparison between (<b>a</b>) experimental and (<b>b</b>) numerical weld nugget size for uncoated steel. The scale bar shows the temperature in degrees Celcius.</p>
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<p>A representative comparison between (<b>a</b>) experimental and (<b>b</b>) predicted weld nugget size for AlSi coated steel. The scale bar shows the temperature in degrees Celsius.</p>
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<p>A representative comparison between experimental and predicted weld nugget sizes for AlSi coated steel and non-coated (legend: NC) steel.</p>
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<p>A representative comparison between experimental and predicted weld nugget sizes for AlSi-coated steel and non-coated (legend: NC) steel. The welding time axis represents the number of cycles in each of the three pulses, such that the total welding time is found by multiplying by 3.</p>
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12 pages, 3077 KiB  
Article
Dexel-Based Simulation of Directed Energy Deposition Additive Manufacturing
by Volker Böß, Berend Denkena, Marc-André Dittrich, Talash Malek and Sven Friebe
J. Manuf. Mater. Process. 2021, 5(1), 9; https://doi.org/10.3390/jmmp5010009 - 11 Jan 2021
Cited by 6 | Viewed by 4349
Abstract
Additive manufacturing is typically a flexible alternative to conventional manufacturing processes. However, manufacturing costs increase due to the effort required to experimentally determine optimum process parameters for customized products or small batches. Therefore, simulation models are needed in order to reduce the amount [...] Read more.
Additive manufacturing is typically a flexible alternative to conventional manufacturing processes. However, manufacturing costs increase due to the effort required to experimentally determine optimum process parameters for customized products or small batches. Therefore, simulation models are needed in order to reduce the amount of effort necessary for experimental testing. For this purpose, a novel technological simulation method for directed energy deposition additive manufacturing is presented here. The Dexel-based simulation allows modeling of additive manufacturing of varying geometric shapes by considering multi-axis machine tool kinematics and local process conditions. The simulation approach can be combined with the simulation of subtractive processes, which enables integrated digital process chains. Full article
(This article belongs to the Special Issue Advances in Modelling of Machining Operations)
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<p>Dexel-based simulation of a milling process.</p>
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<p>Algebraic operations: (<b>a</b>) initial situation; (<b>b</b>) intersection; (<b>c</b>) union.</p>
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<p>Calculation steps for material deposition and transformation of the virtual tool shape into workpiece. Upper row: Calculation steps for a single Dexel line. Lower row: Sequence of calculation steps for the entire Dexel model. (<b>a</b>) Initial situation; (<b>b</b>) complement of workpiece; (<b>c</b>) subtraction of virtual tool; (<b>d</b>) complement; (<b>e</b>) resulting union of workpiece and deposited material with Dexels.</p>
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<p>Dexel representation.</p>
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<p>Material deposition in directed energy deposition (DED) processes.</p>
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<p>Simulation of two different weld seam geometries with the developed method in IFW CutS. (<b>a</b>) Simulation of direction change of the tool path with square weld seam; (<b>b</b>) two-layer weld seam with geometry variation (infeed and outfeed).</p>
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<p>Measured height and standard deviation for different travel velocity and currents.</p>
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<p>Quadratic regression model for prediction of height and width of deposited material.</p>
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<p>Comparison between the simulated and measured deposited materials (<b>a</b>) from the experiment; (<b>b</b>) measured by the tactile measuring system; (<b>c</b>) simulated in CutS.</p>
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20 pages, 4429 KiB  
Article
Prediction of Shearing and Ploughing Constants in Milling of Inconel 718
by Chi-Jen Lin, Yu-Ting Lui, Yu-Fu Lin, Hsian-Bing Wang, Steven Y. Liang and Jiunn-Jyh Junz Wang
J. Manuf. Mater. Process. 2021, 5(1), 8; https://doi.org/10.3390/jmmp5010008 - 11 Jan 2021
Cited by 8 | Viewed by 3159
Abstract
The present study proposes an integrated prediction model for both shearing and ploughing constants for the peripheral milling of Inconel 718 by using a preidentified mean normal friction coefficient. An equation is presented for the identification of normal mean friction angle of oblique [...] Read more.
The present study proposes an integrated prediction model for both shearing and ploughing constants for the peripheral milling of Inconel 718 by using a preidentified mean normal friction coefficient. An equation is presented for the identification of normal mean friction angle of oblique cutting in milling. A simplified oblique cutting model is adopted for obtaining the shear strain and shearing constants for a tool of given helix angle, radial rake angle, and honed edge radius. The shearing and ploughing constants predicted from analytical model using the Merchant’s shear angle formula and the shear flow stress from the selected Johnson–Cook material law are shown to be consistent with the experimental results. The experimentally identified normal friction angles and shearing and edge ploughing constants for the Inconel 718 milling process are demonstrated to have approximately constant values irrespective of the average chip thickness. Moreover, the predicted forces obtained in milling aged Inconel 718 alloy are in good agreement with the experimental force measurements reported in the literature. Without considering the thermal–mechanical coupling effect in the material law, the presented model is demonstrated to work well for milling of both annealed and aged Inconel 718. Full article
(This article belongs to the Special Issue Advances in Modelling of Machining Operations)
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<p>Schematic representation of the peripheral milling process.</p>
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<p>Force components in Merchant circle [<a href="#B1-jmmp-05-00008" class="html-bibr">1</a>].</p>
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<p>Tool edge ploughing mechanism of orthogonal cutting.</p>
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<p>Oblique cutting geometry.</p>
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<p>(<b>a</b>) Verification experiment setup; (<b>b</b>) Inconel 718 workpiece; (<b>c</b>) end mill.</p>
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<p>Comparison of experimental and predicted results for side tool edge cutting constants as a function of average chip thickness. Note that (<b>a</b>–<b>c</b>) show the shearing constants, while (<b>d</b>–<b>f</b>) show the edge ploughing constants.</p>
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<p>Comparison of experimental and predicted results for side tool edge cutting constants as a function of average chip thickness. Note that (<b>a</b>–<b>c</b>) show the shearing constants, while (<b>d</b>–<b>f</b>) show the edge ploughing constants.</p>
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<p>Experimental results for lumped cutting constants (LGCCs) as a function of average chip thickness. (<b>a</b>–<b>c</b>) are for the tangential, radial and axial cutting constants, respectively.</p>
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<p>Experimental and predicted average side forces and percentage errors (<b>a</b>) for three sets of (ai, bi) at <span class="html-italic">a<sub>e</sub></span> = 1.2 mm, and (<b>b</b>) for three sets of (ci, di) at <span class="html-italic">a<sub>e</sub></span> = 1.5 mm.</p>
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<p>Comparison of predicted (lines) and experimental (symbols) milling forces. Note that figures (<b>a</b>–1), (<b>b</b>–1), and (<b>c</b>–1) refer to cutting speeds of 40, 60, and 80 m/min, respectively from [<a href="#B29-jmmp-05-00008" class="html-bibr">29</a>], while figures (<b>a</b>–2), (<b>b</b>–2), and (<b>c</b>–2) refer to predictions by the presented model at two different tool edge radius values along with traced experimental forces. Note that forces in <span class="html-italic">X</span>- and <span class="html-italic">Y</span>-directions on the right are switched for direct comparison with results from [<a href="#B29-jmmp-05-00008" class="html-bibr">29</a>].</p>
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23 pages, 3622 KiB  
Review
Kinematic Fields Measurement during Orthogonal Cutting Using Digital Images Correlation: A Review
by Haythem Zouabi, Madalina Calamaz, Vincent Wagner, Olivier Cahuc and Gilles Dessein
J. Manuf. Mater. Process. 2021, 5(1), 7; https://doi.org/10.3390/jmmp5010007 - 9 Jan 2021
Cited by 11 | Viewed by 3047
Abstract
Understanding the mechanism of chip formation during orthogonal cutting requires a local measurement of the displacement and strain fields in the cutting zone. These measurements can then be used in order to enhance/validate numerical simulation of metal cutting or calibrate material behavior laws [...] Read more.
Understanding the mechanism of chip formation during orthogonal cutting requires a local measurement of the displacement and strain fields in the cutting zone. These measurements can then be used in order to enhance/validate numerical simulation of metal cutting or calibrate material behavior laws for a better prediction of the thermomechanical loads inside the cutting zone. Particle tracking to identify the strain localization that is exhibited in the Adiabatic Shear Band (ASB) is a challenging task. These local measurements can be determined by images post-processing while using the Digital Image Correlation (DIC) technique or analytical models using streamline models or by micro grid analysis. Recently, the use of the DIC technique is widely increased. Texture quality has been shown to be an important factor. Various techniques of surface preparation are then discussed and classified in terms of the created pattern size. Tools for texture analysis are presented. The technique suitability for the kinematic fields measurement while using the DIC technique during machining is discussed. Various optical systems of the literature employed in the context of kinematic fields measurement during machining are discussed in this paper. The recent advances on the design of optical systems are given. Finally, the results of kinematic fields measurement during machining metallic alloys are analyzed. Full article
(This article belongs to the Special Issue Optimization and Simulation of Solid State Manufacturing Processes)
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<p>Cutting zones and the region of interest.</p>
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<p>Synoptic scheme of the article.</p>
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<p>Schematic illustration of the principle of the Digital Image Correlation (DIC) with local approach.</p>
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<p>Illustration of the displacement vector of the speckle pattern center according to [<a href="#B51-jmmp-05-00007" class="html-bibr">51</a>].</p>
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<p>Surface preparation: techniques and texture assessment tools.</p>
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<p>Occurence of residual displacements over 5 successives images [<a href="#B43-jmmp-05-00007" class="html-bibr">43</a>].</p>
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<p>Scheme of the optical system used in [<a href="#B29-jmmp-05-00007" class="html-bibr">29</a>] and described in [<a href="#B46-jmmp-05-00007" class="html-bibr">46</a>].</p>
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<p>Utra-high speed imaging system developed by [<a href="#B38-jmmp-05-00007" class="html-bibr">38</a>]: (<b>a</b>) four synchronized high-speed camera and (<b>b</b>) four combined pulsed laser sources.</p>
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<p>Optical system used in [<a href="#B24-jmmp-05-00007" class="html-bibr">24</a>] to study the surface integrity.</p>
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<p>MADMACS dual-spectrum optical system developed by [<a href="#B68-jmmp-05-00007" class="html-bibr">68</a>].</p>
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<p>Visible and infrared imaging system (VISIR) developed by [<a href="#B49-jmmp-05-00007" class="html-bibr">49</a>].</p>
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<p>(<b>a</b>) Displacement field in the primary shear zone for automatic measurement of the shear angle and (<b>b</b>) displacement field in the subsurface of the workpiece during orthogonal cutting [<a href="#B24-jmmp-05-00007" class="html-bibr">24</a>].</p>
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<p>(<b>a</b>) Initial and final position of the three investigated points (P1, P2, and P3), (<b>b</b>) strain rate evolution for each point over the sequance of images [<a href="#B43-jmmp-05-00007" class="html-bibr">43</a>].</p>
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<p>(<b>a</b>) Captured visual images at different instants during one segment chip formation; (<b>b</b>) crack propagation within the PSZ; and, (<b>c</b>) cumulative Henky strain field and (<b>d</b>) the equivalent strain rate field [<a href="#B49-jmmp-05-00007" class="html-bibr">49</a>].</p>
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<p>Kinematic fields measurement: analytical technique and the DIC technique.</p>
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25 pages, 10575 KiB  
Article
High-Temperature Equal-Channel Angular Pressing of a T6-Al-Cu-Li-Mg-Ag-Zr-Sc Alloy
by Marcello Cabibbo and Chiara Paoletti
J. Manuf. Mater. Process. 2021, 5(1), 6; https://doi.org/10.3390/jmmp5010006 - 5 Jan 2021
Cited by 1 | Viewed by 2641
Abstract
Equal-channel angular pressing (ECAP) is known to induce significant grain refinement and formation of tangled dislocations within the grains. These are induced to evolve to form low-angle boundaries (i.e., cell boundaries) and eventually high-angle boundaries (i.e., grain boundaries). On the other hand, the [...] Read more.
Equal-channel angular pressing (ECAP) is known to induce significant grain refinement and formation of tangled dislocations within the grains. These are induced to evolve to form low-angle boundaries (i.e., cell boundaries) and eventually high-angle boundaries (i.e., grain boundaries). On the other hand, the precipitation sequence of age hardening aluminum alloys can be significantly affected by pre-straining and severe plastic deformation. Thus, ECAP is expected to influence the T6 response of aluminum alloys. In this study, a complex Al-Cu-Mg-Li-Ag-Zr-Sc alloy was subjected to ECAP following different straining paths. The alloy was ECAP at 460 K via route A, C, and by forward-backward route A (FB-route A) up to four passes. That is, ECAP was carried out imposing billet rotation between passes (route A), billet rotation by +90° between passes (route C), and billet rotation by +90° and inversion upside down between passes (FB-route A). The alloy was also aged at 460 K for different durations after ECAP. TEM microstructure inspections showed a marked influence of the different shearing deformations induced by ECAP on the alloy aging response. The precipitation kinetics of the different hardening secondary phases were affected by shearing deformation and tangled dislocations. In particular, the T1-Al2CuLi phase was the one that mostly showed a precipitation sequence speed up induced by the tangled dislocations formed during ECAP. The T1 phase was found to grow with aging time according to the Lifshitz-Slyozov-Wagner low-power regime. Full article
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<p>(<b>a</b>) Characteristic deformation directions in the Equal-Channel Angular Pressed (ECAP) billet. The <span class="html-italic">Y</span>-plane contains the shear deformation bands induced by ECAP and it is the one selected for TEM inspections. ECAP direction is along the billet <span class="html-italic">X</span> direction; (<b>b</b>) scheme of the activated shear planes in routes: A, FB-A, C.</p>
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<p>Hardness, <span class="html-italic">H</span>, vs. aging time, <span class="html-italic">t</span>, of the alloy annealing at 813 K/4 h and aging at 460 K, (<b>a</b>), BF-TEM showing the microstructure at T6 hardness peak condition (aging at 460 K/6 h), [002]<sub>Al</sub>, (<b>b</b>), [111]<sub>Al</sub>, (<b>c</b>), and [210]<sub>Al</sub>, (<b>d</b>). Polarized optical micrograph showing the alloy T6 grained structure, (<b>e</b>).</p>
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<p>Al<sub>3</sub>(Sc<sub>1−x</sub>,Zr<sub>x</sub>) dispersoids showing Ashby-Brown strain contrast in the alloy tempered at 813 K/4 h + aging at 460 K/6 h. Inset is a detail of two <span class="html-italic">β′</span> dispersoids (dark rounded particles) showing an Ashby-Brown contrast.</p>
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<p>BF-TEM of ECAP-A/4, where the crystal is oriented to show (100)<sub>Al</sub>, (<b>a</b>), (110)<sub>Al</sub>, (<b>b</b>), and (111)<sub>Al</sub>, (<b>c</b>); ECAP-C/4 along (100)<sub>Al</sub>, (<b>d</b>), (110)<sub>Al</sub>, (<b>e</b>), and (111)<sub>Al</sub>, (<b>f</b>); FB-ECAP-A/4 along (100)<sub>Al</sub>, (<b>g</b>), (110)<sub>Al</sub>, (<b>h</b>), and (111)<sub>Al</sub>, (<b>i</b>). Al-Cu clusters and GP zones were detected along the three planes; tiny needle-like <span class="html-italic">θ</span>-Al<sub>2</sub>Cu phase particles were detected along [100]<sub>Al</sub>; coarser needle-like <span class="html-italic">T<sub>1</sub></span>-Al<sub>2</sub>CuLi phase particles were detected along [110]<sub>Al</sub>; rounded and somewhat quasi-equiaxed <span class="html-italic">δ</span>-Al<sub>3</sub>Li phase particles were detected along [111]<sub>Al</sub>.</p>
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<p>BF-TEM of ECAP-A/4, where the crystal is oriented to show (100)<sub>Al</sub>, (<b>a</b>), (110)<sub>Al</sub>, (<b>b</b>), and (111)<sub>Al</sub>, (<b>c</b>); ECAP-C/4 along (100)<sub>Al</sub>, (<b>d</b>), (110)<sub>Al</sub>, (<b>e</b>), and (111)<sub>Al</sub>, (<b>f</b>); FB-ECAP-A/4 along (100)<sub>Al</sub>, (<b>g</b>), (110)<sub>Al</sub>, (<b>h</b>), and (111)<sub>Al</sub>, (<b>i</b>). Al-Cu clusters and GP zones were detected along the three planes; tiny needle-like <span class="html-italic">θ</span>-Al<sub>2</sub>Cu phase particles were detected along [100]<sub>Al</sub>; coarser needle-like <span class="html-italic">T<sub>1</sub></span>-Al<sub>2</sub>CuLi phase particles were detected along [110]<sub>Al</sub>; rounded and somewhat quasi-equiaxed <span class="html-italic">δ</span>-Al<sub>3</sub>Li phase particles were detected along [111]<sub>Al</sub>.</p>
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<p>LM BF-TEM showing the alloy grain structure along [110]<sub>Al</sub> after ECAP-A/4, (<b>a</b>), ECAP-C/4, (<b>b</b>), and FB-ECAP-A/4, (<b>c</b>). Insets are Selected Area Diffraction Patterns (SAEDPs) showing the process of grain refinement through formation of concentric diffraction rings.</p>
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<p>Tangled dislocation density, <span class="html-italic">ρ<sub>disl</sub></span>, cell, <span class="html-italic">D<sub>cell</sub></span>, and grain, <span class="html-italic">D<sub>g</sub></span>, size, after ECAP-A/4, ECAP-C/4, and FB-ECAP-A/4 at 460 K, (<b>a</b>); measured mean number density of GP zones and precipitates, <span class="html-italic">n<sub>V</sub></span>, of all detected secondary phases <span class="html-italic">T</span><sub>1</sub>, <span class="html-italic">θ</span>, and <span class="html-italic">δ</span> induced to form under ECAP-A/4, ECAP-C/4, and FB-ECAP-A/4 at 460 K, (<b>b</b>); mean longer edge lengths of the <span class="html-italic">T</span><sub>1</sub> platelets, <span class="html-italic">l<sub>T</sub></span><sub>1</sub>, at alloy T6-condition and after ECAP at 460 K are also reported, (<b>c</b>).</p>
Full article ">Figure 6 Cont.
<p>Tangled dislocation density, <span class="html-italic">ρ<sub>disl</sub></span>, cell, <span class="html-italic">D<sub>cell</sub></span>, and grain, <span class="html-italic">D<sub>g</sub></span>, size, after ECAP-A/4, ECAP-C/4, and FB-ECAP-A/4 at 460 K, (<b>a</b>); measured mean number density of GP zones and precipitates, <span class="html-italic">n<sub>V</sub></span>, of all detected secondary phases <span class="html-italic">T</span><sub>1</sub>, <span class="html-italic">θ</span>, and <span class="html-italic">δ</span> induced to form under ECAP-A/4, ECAP-C/4, and FB-ECAP-A/4 at 460 K, (<b>b</b>); mean longer edge lengths of the <span class="html-italic">T</span><sub>1</sub> platelets, <span class="html-italic">l<sub>T</sub></span><sub>1</sub>, at alloy T6-condition and after ECAP at 460 K are also reported, (<b>c</b>).</p>
Full article ">Figure 7
<p>Microstructures of the post-ECAP 460 K/6 h annealed alloys, showing the different secondary-phase precipitates. BF-TEM of ECAP-A/4 along (100)<sub>Al</sub>, (<b>a</b>), (110)<sub>Al</sub>, (<b>b</b>), and (111)<sub>Al</sub> planes, (<b>c</b>); ECAP-C/4 along (100)<sub>Al</sub>, (<b>d</b>), (110)<sub>Al</sub>, (<b>e</b>), and (111)<sub>Al</sub> planes, (<b>f</b>); FB-ECAP-A/4 along (100)<sub>Al</sub>, (<b>g</b>), (110)<sub>Al</sub>, (<b>h</b>), and (111)<sub>Al</sub> planes, (<b>i</b>).</p>
Full article ">Figure 7 Cont.
<p>Microstructures of the post-ECAP 460 K/6 h annealed alloys, showing the different secondary-phase precipitates. BF-TEM of ECAP-A/4 along (100)<sub>Al</sub>, (<b>a</b>), (110)<sub>Al</sub>, (<b>b</b>), and (111)<sub>Al</sub> planes, (<b>c</b>); ECAP-C/4 along (100)<sub>Al</sub>, (<b>d</b>), (110)<sub>Al</sub>, (<b>e</b>), and (111)<sub>Al</sub> planes, (<b>f</b>); FB-ECAP-A/4 along (100)<sub>Al</sub>, (<b>g</b>), (110)<sub>Al</sub>, (<b>h</b>), and (111)<sub>Al</sub> planes, (<b>i</b>).</p>
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<p>Secondary-phase precipitates identified by SAEDPs: <span class="html-italic">T</span><sub>1</sub>, (<b>a</b>), <span class="html-italic">Ω</span> and <span class="html-italic">T</span><sub>1</sub>, (<b>b</b>), <span class="html-italic">δ</span> and <span class="html-italic">θ</span>, (<b>c</b>), <span class="html-italic">S</span> and <span class="html-italic">T</span><sub>1</sub>, (<b>d</b>).</p>
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<p>Tangled dislocation density, <span class="html-italic">ρ<sub>disl</sub></span>, cell, <span class="html-italic">D<sub>cell</sub></span>, and grain, <span class="html-italic">D<sub>g</sub></span>, size, (<b>a</b>); secondary-phase precipitate mean number density, <span class="html-italic">n<sub>V</sub></span>(<span class="html-italic">T</span><sub>1</sub>), <span class="html-italic">n<sub>V</sub></span>(<span class="html-italic">Ω</span>), <span class="html-italic">n<sub>V</sub></span>(<span class="html-italic">θ</span>), <span class="html-italic">n<sub>V</sub></span>(<span class="html-italic">δ</span>), and <span class="html-italic">n<sub>V</sub></span>(<span class="html-italic">S</span>), (<b>b</b>); and size, <span class="html-italic">d<sub>T</sub></span><sub>1</sub>, <span class="html-italic">d<sub>Ω</sub></span>, <span class="html-italic">d<sub>θ</sub></span>, <span class="html-italic">d<sub>δ</sub></span>, and <span class="html-italic">d<sub>S</sub></span> (<b>c</b>) formed after ECAP-A/4, ECAP-C/4, and FB-ECAP-A/4 and subsequent aging at 460 K/6 h (overaged T8 condition). Mean longer edge length of the <span class="html-italic">T</span><sub>1</sub> platelets, <span class="html-italic">l<sub>T</sub></span><sub>1</sub>, is also reported, (<b>c</b>).</p>
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<p><span class="html-italic">δ</span>′/<span class="html-italic">S</span>′ precipitate cutting from shearing deformation induced by ECAP, (<b>a</b>), and the helicoid dislocation pinning phenomenon occurring at <span class="html-italic">δ</span>′/<span class="html-italic">S</span>′ precipitates, (<b>b</b>). TEM micrographs refer to FB-ECAP-A/4 + aging at 460 K/2 h (T8 hardness peak condition).</p>
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<p>Aging curves after ECAP-A/4, ECAP-C/4, and FB-ECAP-A/4 and subsequent aging at 540 K for 10, 20, 30, 60, and 90 min, 2, 4, 6, 8, 10, and 16 h, and 1 and 2 days (same times of the undeformed alloy T6-heat treatment). Experimental errors are essentially within the datapoint.</p>
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<p>Microstructures of the post-ECAP 460 K aged alloys, with aging duration set at the maximum alloy hardness peak obtained after ECAP and subsequent aging at 460 K (3 h for ECAP-A/4, 5 h for ECAP-C/4, and 2 h for FB-ECAP-A/4). BF-TEM of ECAP-A/4 at [002]<sub>Al</sub>, (<b>a</b>), [111]<sub>Al</sub>, (<b>b</b>), and [210]<sub>Al</sub>, (<b>c</b>); ECAP-C/4 at [002]<sub>Al</sub>, (<b>d</b>), [111]<sub>Al</sub>, (<b>e</b>), and [210]<sub>Al</sub>, (<b>f</b>); FB-ECAP-A/4 at [002]<sub>Al</sub>, (<b>g</b>), [111]<sub>Al</sub>, (<b>h</b>), and [210]<sub>Al</sub>, (<b>i</b>).</p>
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<p>Tangled dislocation density, <span class="html-italic">ρ<sub>disl</sub></span>, cell, <span class="html-italic">D<sub>cell</sub></span>, and grain, <span class="html-italic">D<sub>g</sub></span>, size (<b>a</b>), secondary-phase precipitate mean number density, <span class="html-italic">n<sub>V</sub></span>(<span class="html-italic">T</span><sub>1</sub>), <span class="html-italic">n<sub>V</sub></span>(<span class="html-italic">Ω</span>), <span class="html-italic">n<sub>V</sub></span>(<span class="html-italic">θ</span>), <span class="html-italic">n<sub>V</sub></span>(<span class="html-italic">δ</span>), <span class="html-italic">n<sub>V</sub></span>(<span class="html-italic">S</span>) (<b>b</b>), and size, <span class="html-italic">d<sub>T</sub></span><sub>1</sub>, <span class="html-italic">d<sub>Ω</sub></span>, <span class="html-italic">d<sub>θ</sub></span>, <span class="html-italic">d<sub>δ</sub></span>, <span class="html-italic">d<sub>S</sub></span> (<b>c</b>) formed after ECAP-A/4, ECAP-C/4, and FB-ECAP-A/4 and subsequent aging at 460 K to reach alloy hardness peak (3 h, for ECAP-A/4, 5 h, for ECAP-C/4, and 2 h, for FB-ECAP-A/4). Mean longer edge lengths of the <span class="html-italic">T</span><sub>1</sub> platelets, <span class="html-italic">l<sub>T</sub></span><sub>1</sub>, are also reported in (<b>c</b>).</p>
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<p>Relative frequency of the strengthening secondary-phase precipitates in T6 undeformed alloy, after ECAP-A/4, ECAP-C/4, and FB-ECAP-A/4 at 460 K, after ECAP-A/4 + aging at 460 K/3 h, ECAP-C/4 + aging at 460 K/5 h, FB-ECAP-A/4 + aging at 460 K/2 h, and after ECAP-A/4, ECAP-C/4, and FB-ECAP-A/4 + aging at 460 K/6 h. In the histogram, <span class="html-italic">E</span> stands for ECAP.</p>
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<p>T1 + Ω precipitate power-law growth. T1 + Ω plate/lath long-edge, <span class="html-italic">l<sub>T1 + Ω</sub></span>, vs. <span class="html-italic">t</span><sup>1/3</sup>, where <span class="html-italic">t</span> is the aging time, (<b>a</b>), tangled dislocation density, <span class="html-italic">ρ<sub>disl</sub></span>, vs. number density of the T1 + Ω precipitates, <span class="html-italic">n<sub>V</sub></span>(<span class="html-italic">T</span><sub>1</sub> + <span class="html-italic">Ω</span>), (<b>b</b>). Data refer to ECAP-A/4, ECAP-C/4, and FB-ECAP-A/4 at 460 K, ECAP-A/4 + aging at 460 K/3 h, ECAP-C/4 + aging at 460 K/5 h, FB-ECAP-A/4 + aging at 460 K/2 h, and ECAP-A/4, ECAP-C/4, and FB-ECAP-A/4 + aging at 460 K/6 h.</p>
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17 pages, 6488 KiB  
Article
Optimization of Thin Walls with Sharp Corners in SS316L and IN718 Alloys Manufactured with Laser Metal Deposition
by Juan Carlos Pereira, Herman Borovkov, Fidel Zubiri, Mari Carmen Guerra and Josu Caminos
J. Manuf. Mater. Process. 2021, 5(1), 5; https://doi.org/10.3390/jmmp5010005 - 5 Jan 2021
Cited by 12 | Viewed by 4064
Abstract
In this work, the manufacture of thin walls with sharp corners has been optimized by adjusting the limits of a 3-axis cartesian kinematics through data recorded and analyzed off-line, such as axis speed, acceleration and the positioning of the X and Y axes. [...] Read more.
In this work, the manufacture of thin walls with sharp corners has been optimized by adjusting the limits of a 3-axis cartesian kinematics through data recorded and analyzed off-line, such as axis speed, acceleration and the positioning of the X and Y axes. The study was carried out with two powder materials (SS316L and IN718) using the directed energy deposition process with laser. Thin walls were obtained with 1 mm thickness and only one bead per layer and straight/sharp corners at 90°. After adjusting the in-position parameter G502 for positioning precision on the FAGOR 8070 CNC system, it has been possible to obtain walls with minimal accumulation of material in the corner, and with practically constant layer thickness and height, with a radii of internal curvature between 0.11 and 0.24 mm for two different precision configuration. The best results have been obtained by identifying the correct balance between the decrease in programmed speed and the precision in the positioning to reach the point defined as wall corner, with speed reductions of 29% for a programmed speed of 20 mm/s and 61% for a speed of 40 mm/s. The walls show minimal defects such as residual porosities, and the microstructure is adequate. Full article
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Figure 1

Figure 1
<p>Schematic representation of coaxial LMD process with powder (LMD-p) [<a href="#B3-jmmp-05-00005" class="html-bibr">3</a>] (With permission of TRUMPF Laser- und Systemtechnik GmbH).</p>
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<p>Characterization of pre-alloyed gas atomized powders. Morphology of the particles in SEM image (20 kV), and Granulometry and distribution analysis. SS316L powder (<b>left</b>) and IN718 Powder (<b>right</b>).</p>
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<p>Characterization of pre-alloyed gas atomized powders. Morphology of the particles in SEM image (20 kV), and Granulometry and distribution analysis. SS316L powder (<b>left</b>) and IN718 Powder (<b>right</b>).</p>
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<p>Cartesian kinematics LMD station in LORTEK. Detail of the station (<b>left</b>) and detail of the Coax40 nozzle with the powder flow (<b>right</b>).</p>
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<p>Thin walls manufactured by LMD-powder. 20 × 40 × 25 mm wall made of SS316L without process control and different “e” parameters (<b>left</b>) and 30 × 50 × 10 mm wall made of IN718 with process control (<b>right</b>).</p>
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<p>Deviation of the deposition path when approaching a straight corner in the LMD process.</p>
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<p>Position register (X and Y axes) in the vicinity of the corner for different values of e: X.XX and two speeds (v = 20 mm/s and v = 40 mm/s).</p>
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<p>Speed recording as a function of position on each axis in the vicinity of the corner. v = 20 mm/s (<b>left</b>) and v = 40 mm/s (<b>right</b>).</p>
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<p>3D plot of position as a function of velocity (precision e = 0.4) for two successive layers.</p>
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<p>3D plot of position as a function of velocity (precision e = 0.4) for two successive layers.</p>
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<p>Radius of curvature in thin walls made with powder LMD. SS316L walls without power control with e = 0.40 (<b>left</b>), e = 0.04 (<b>center</b>) and wall in IN718 with e = 0.40 and power control (<b>right</b>).</p>
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<p>Photomicrograph of walls in a polished state. SS316L walls of sample 1 (<b>left</b>), sample 2 (<b>center</b>) and sample 3 wall manufactured with IN718 (<b>right</b>).</p>
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<p>Photomicrographs by LOM of the walls made with SS316L. Sample 1 (e = 0.40, <b>left</b>) and Sample 2 (e = 0.04, <b>right</b>). Different study areas: Interface with the substrate (200× below), at 5 mm wall height (500× center) and 10 mm wall height (500× above).</p>
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<p>Photomicrographs by LOM of sample 3 (IN718, e = 0.40 and power control) in different areas. Interface with substrate (200× bottom), at 5 mm wall height (500× center) and 10 mm wall height (500× top).</p>
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<p>Microhardness profile on the walls made of SS316L (samples 1 and 2) and the microstructure in each area of the walls.</p>
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<p>Microhardness profile on the wall manufactured with IN718 (sample 3) and the microstructure in each area of the wall.</p>
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15 pages, 3008 KiB  
Article
Potentials of Vitrified and Elastic Bonded Fine Grinding Worms in Continuous Generating Gear Grinding
by Maximilian Schrank, Jens Brimmers and Thomas Bergs
J. Manuf. Mater. Process. 2021, 5(1), 4; https://doi.org/10.3390/jmmp5010004 - 5 Jan 2021
Cited by 4 | Viewed by 2841
Abstract
Continuous generating gear grinding with vitrified grinding worms is an established process for the hard finishing of gears for high-performance transmissions. Due to the increasing requirements for gears in terms of power density, the required surface roughness is continuously decreasing. In order to [...] Read more.
Continuous generating gear grinding with vitrified grinding worms is an established process for the hard finishing of gears for high-performance transmissions. Due to the increasing requirements for gears in terms of power density, the required surface roughness is continuously decreasing. In order to meet the required tooth flank roughness, common manufacturing processes are polish grinding with elastic bonded grinding tools and fine grinding with vitrified grinding tools. The process behavior and potential of the different bonds for producing super fine surfaces in generating gear grinding have not been sufficiently scientifically investigated yet. Therefore, the objective of this report is to evaluate these potentials. Part of the investigations are the generating gear grinding process with elastic bonded, as well as vitrified grinding worms with comparable grit sizes. The potential of the different tool specifications is empirically investigated independent of the grain size, focusing on the influence of the bond. One result of the investigations was that the tooth flank roughness could be reduced to nearly the same values with the polish and the fine grinding tool. Furthermore, a dependence of the roughness on the selected grinding parameters could not be determined. However, it was found out that the profile line after polish grinding is significantly dependent on the process strategy used. Full article
(This article belongs to the Special Issue Progress in Precision Machining)
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Figure 1

Figure 1
<p>Increase of the load-carrying capacity by reducing the roughness according to [<a href="#B13-jmmp-05-00004" class="html-bibr">13</a>].</p>
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<p>Experimental setup and procedure.</p>
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<p>Influence of the waviness filter and roughness profile of polish and fine ground flanks.</p>
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<p>Results of pre-grinding as well as polish and fine grinding with a feed variation.</p>
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<p>Results of pre-grinding with different dressing tool technologies.</p>
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<p>The results of polish ground gears with different relative infeed.</p>
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<p>Results of polish and fine grinding with a two-flank profile dressing tool.</p>
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<p>Results of polish and fine grinding with varied cutting speeds.</p>
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<p>Roughness profiles of pre-, polish, and fine ground gears.</p>
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16 pages, 8374 KiB  
Article
Experimental Investigation of Dimensional Precision of Deep Drawn Cups Using Direct Polymer Additive Tooling
by Georg Bergweiler, Falko Fiedler, Ahsan Shaukat and Bernd Löffler
J. Manuf. Mater. Process. 2021, 5(1), 3; https://doi.org/10.3390/jmmp5010003 - 30 Dec 2020
Cited by 11 | Viewed by 3662
Abstract
While deep drawing of sheet metals is economical at high volumes, it can be very costly for manufacturing prototypes, mainly due to high tooling costs. Additively manufactured polymer tools have the potential to be more cost-efficient for small series, but they are softer [...] Read more.
While deep drawing of sheet metals is economical at high volumes, it can be very costly for manufacturing prototypes, mainly due to high tooling costs. Additively manufactured polymer tools have the potential to be more cost-efficient for small series, but they are softer and thus less resilient than conventional steel tools. This work aimed to study the dimensional precision of such tools using a standard cup geometry. Tools were printed with FFF using two different materials, PLA and CF-PA. A test series of 20 parts was drawn from each tool. Afterwards, the dimensional precisions of the drawn parts were analyzed using an optical measuring system. The achieved dimensional accuracy of the first drawn cup using the PLA toolset was 1.98 mm, which was further improved to 1.04 mm by altering shrinkage and springback allowances. The repeatability of the deep drawing process for the CF-reinforced PA tool was 0.17 mm during 20 drawing operations and better than that of the PLA tool (1.17 mm). To conclude, deep drawing of standard cups is doable using direct polymer additive tooling with a dimensional accuracy of 1.04 mm, which can be further improved by refining allowances incorporated to the CAD model being printed. Full article
(This article belongs to the Special Issue Advances in Sheet Metal Forming and Structures)
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Figure 1

Figure 1
<p>Working principle of deep drawing: (<b>a</b>) a sheet metal blank is inserted between die and blank holder; (<b>b</b>) forces on both the blank holder and the punch are applied; (<b>c</b>) final cup geometry including a flange is formed.</p>
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<p>Working principle of FFF process [<a href="#B16-jmmp-05-00003" class="html-bibr">16</a>] (reproduced with the permission of the Verein Deutscher Ingenieure e.V.).</p>
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<p>Properties of different AM polymers (for abbreviations and chemical formulae, see <a href="#app1-jmmp-05-00003" class="html-app">Appendix A</a>): (<b>a</b>) tensile strength plotted over flexural strength; (<b>b</b>) cost per kilogram plotted over density [<a href="#B17-jmmp-05-00003" class="html-bibr">17</a>,<a href="#B18-jmmp-05-00003" class="html-bibr">18</a>,<a href="#B19-jmmp-05-00003" class="html-bibr">19</a>,<a href="#B20-jmmp-05-00003" class="html-bibr">20</a>,<a href="#B21-jmmp-05-00003" class="html-bibr">21</a>,<a href="#B22-jmmp-05-00003" class="html-bibr">22</a>,<a href="#B23-jmmp-05-00003" class="html-bibr">23</a>,<a href="#B24-jmmp-05-00003" class="html-bibr">24</a>,<a href="#B25-jmmp-05-00003" class="html-bibr">25</a>,<a href="#B26-jmmp-05-00003" class="html-bibr">26</a>].</p>
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<p>Schematic of deep drawing parameters.</p>
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<p>Isometric view of the tools including mounting holes: (<b>a</b>) punch; (<b>b</b>) blank holder; (<b>c</b>) die.</p>
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<p>Surface scans of CF-PA material showing the stair stepping effect at a radius of 3 mm: (<b>a</b>) vertical view; (<b>b</b>) isometric view.</p>
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<p>Measuring points of tool and cup surface: (<b>a</b>) measuring points distribution for cup; (<b>b</b>) distribution of eight sets of points for average (here exemplary on cup surface); (<b>c</b>) measuring points distribution for punch; (<b>d</b>) measuring points distribution for die.</p>
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<p>Scanned outer surface of a cup compared to its reference geometry (the comparison to the CAD geometry is exaggerated for easier understanding).</p>
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<p>Deviation values of PLA tools and their drawn cups compared to their CAD reference geometries: (<b>a</b>) pillar chart of the zeroth drawn tool and first drawn cup; (<b>b</b>) color map of the outer surface of first drawn cup; (<b>c</b>) color map of the outer surface of the 20th drawn cup; (<b>d</b>) pillar chart of the 15th drawn tool and 20th drawn cup.</p>
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<p>Deviation values of CF-reinforced PA tools and their drawn cups compared to their CAD reference geometries (please note the different scale of the pillar chart compared to <a href="#jmmp-05-00003-f009" class="html-fig">Figure 9</a>): (<b>a</b>) pillar chart of zeroth drawn tool and first drawn cup; (<b>b</b>) color map of the outer surface of the first drawn cup; (<b>c</b>) color map of the outer surface of the 20th drawn cup; (<b>d</b>) pillar chart of the 15th drawn tool and 20th drawn cup.</p>
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<p>Color map of the outer surface of 20th drawn cup compared to the first drawn cup: (<b>a</b>) color map of cup drawn from CF-reinforced PA tools; (<b>b</b>) color map of cup drawn from PLA tools.</p>
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<p>Deviation values of different measuring points of cups during 20 drawing operations compared to the CAD reference geometry: (<b>a</b>) deviation values of cups drawn from CF-reinforced PA tools; (<b>b</b>) deviation values of cups drawn from PLA tools.</p>
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<p>Color map of deviations of the 20th drawn punch compared to the zeroth drawn punch: (<b>a</b>) CF-PA punch; (<b>b</b>) PLA punch.</p>
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<p>Course of deviation values of different measuring points of the punch during 20 drawing operations compared to the CAD reference geometry: (<b>a</b>) deviation values of CF-reinforced PA punch; (<b>b</b>) deviation values of PLA punch.</p>
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<p>Deviation values of both trials of PLA tools and their drawn cups compared to their CAD reference geometries (please note the different scale of the pillar chart compared to <a href="#jmmp-05-00003-f010" class="html-fig">Figure 10</a>): (<b>a</b>) pillar chart of PLA_001 zeroth drawn tool and first drawn cup and color map of first drawn cup; (<b>b</b>) pillar chart of PLA_001 15th drawn tool and 20th drawn cup and color map of 20th drawn cup; (<b>c</b>) pillar chart of the PLA_002 zeroth drawn tool and first drawn cup and color map of first drawn cup; (<b>d</b>) pillar chart of the PLA_002 18th drawn tool and 20th drawn cup and color map of the 20th drawn cup.</p>
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<p>Color map of deviations of 20th drawn cup compared to the first drawn cup for both PLA trials: (<b>a</b>) cup drawn with PLA_001 tool; (<b>b</b>) cup drawn with PLA_002 with allowances incorporated.</p>
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<p>Course of deviation values of different measuring points of the cup during 20 drawing operations compared to the CAD reference geometry: (<b>a</b>) deviation values of cups drawn with PLA_001 tool; (<b>b</b>) deviation values of cup drawn with PLA_002 tool.</p>
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<p>Color map of deviations of the 20th drawn punch compared to the zeroth drawn punch: (<b>a</b>) PLA_001 punch; (<b>b</b>) PLA_002 punch.</p>
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<p>Course of deviation values of different measuring points of punch during 20 drawing operations compared to the CAD reference geometry: (<b>a</b>) deviation values of PLA_001 punch; (<b>b</b>) deviation values of PLA_002 punch.</p>
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12 pages, 4630 KiB  
Article
Optimization and Tuning of Passive Tuned Mass Damper Embedded in Milling Tool for Chatter Mitigation
by Wenshuo Ma, Jingjun Yu, Yiqing Yang and Yunfei Wang
J. Manuf. Mater. Process. 2021, 5(1), 2; https://doi.org/10.3390/jmmp5010002 - 25 Dec 2020
Cited by 11 | Viewed by 3538
Abstract
Milling tools with a large length–diameter ratio are widely applied in machining structural features with deep depth. However, their high dynamic flexibility gives rise to chatter vibrations, which results in poor surface finish, reduced productivity, and even tool damage. With a passive tuned [...] Read more.
Milling tools with a large length–diameter ratio are widely applied in machining structural features with deep depth. However, their high dynamic flexibility gives rise to chatter vibrations, which results in poor surface finish, reduced productivity, and even tool damage. With a passive tuned mass damper (TMD) embedded inside the arbor, a large length–diameter ratio milling tool with chatter-resistance ability was developed. By modeling the milling tool as a continuous beam, the tool-tip frequency response function (FRF) of the milling tool with TMD was derived using receptance coupling substructure analysis (RCSA), and the gyroscopic effect of the rotating tool was incorporated. The TMD parameters were optimized numerically with the consideration of mounting position based on the maximum cutting stability criterion, followed by the simulation of the effectiveness of the optimized and detuned TMD. With the tool-tip FRF obtained, the chatter stability of the milling process was predicted. Tap tests showed that the TMD was able to increase the minimum real part of the FRF by 79.3%. The stability lobe diagram (SLD) was predicted, and the minimum critical depth of cut in milling operations was enhanced from 0.10 to 0.46 mm. Full article
(This article belongs to the Special Issue Machine Tool Dynamics)
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<p>The milling tool embedded with a tuned mass damper (TMD). 1—end cover; 2—damping element; 3 and 8—gasket; 4—shell; 5—mass block; 6—steel rod; 7—support block.</p>
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<p>Schematic of the coupling subsystems.</p>
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<p>The predicted and experimental tool-tip frequency response functions (FRFs) of the dominant mode.</p>
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<p>The tool-tip FRFs of the idle and rotating milling tool.</p>
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<p>Optimum parameters versus different mass ratios.</p>
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<p>The predicted tool-tip FRFs of the milling tool with TMD optimized by the current study and [<a href="#B16-jmmp-05-00002" class="html-bibr">16</a>].</p>
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<p>The influence of TMD parameter deviation on the negative real part of the tool-tip FRF.</p>
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<p>The experimental setup.</p>
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<p>The predicted and experimental tool-tip FRFs of the milling tool embedded with TMD.</p>
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<p>The SLDs of the milling tool without and with TMD.</p>
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17 pages, 2661 KiB  
Article
Validation of a Coupled Simulation for Machine Tool Dynamics Using a Linear Drive Actuator
by Michael Wiesauer, Christoph Habersohn and Friedrich Bleicher
J. Manuf. Mater. Process. 2021, 5(1), 1; https://doi.org/10.3390/jmmp5010001 - 23 Dec 2020
Cited by 4 | Viewed by 3337
Abstract
In order to ensure high productivity capabilities of machine tools at a low cost but at increased geometric accuracy, modeling of their static and dynamic behavior is a crucial task in structure optimization. The drive control and the frictional forces acting in feed [...] Read more.
In order to ensure high productivity capabilities of machine tools at a low cost but at increased geometric accuracy, modeling of their static and dynamic behavior is a crucial task in structure optimization. The drive control and the frictional forces acting in feed axes significantly determine the machine’s response in the frequency domain. The aim of this study was the accurate modeling and the experimental investigation of dynamic damping effects using a machine tool test rig with three-axis kinematics. For this purpose, an order-reduced finite element model of the mechanical structure was coupled with models of the drive control and of the non-linear friction behavior. In order to validate the individual models, a new actuator system based on a tubular linear drive was used for frequency response measurements during uniaxial carriage movements. A comparison of the dynamic measurements with the simulation results revealed a good match of amplitudes in the frequency domain by considering dynamic damping. Accordingly, the overall dynamic behavior of machine tool structures can be predicted and thus optimized by a coupled simulation at higher level of detail and by considering the damping effects of friction. Dynamic testing with the newly designed actuator is a prerequisite for model validation and control drive parameterization. Full article
(This article belongs to the Special Issue Machine Tool Dynamics)
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<p>Schematic representation of the mechatronic machine tool model.</p>
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<p>Schematic model configuration of the ball screw drive.</p>
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<p>Schematic representation of the cascaded control loop model.</p>
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<p>(<b>a</b>) Measured friction force of the <span class="html-italic">X</span>-axis (spindle pitch 40 mm) during a sinusoidal movement with amplitude 15 mm, feed rate 1000 mm/min and (<b>b</b>) feed rate 3000 mm/min. (<b>c</b>) Comparison of measured/simulated driving forces at feed rate 3000 mm/min.</p>
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<p>(<b>a</b>) Electromagnetic linear motor in tubular form for vibration excitation, (<b>b</b>) associated control case and (<b>c</b>) the measurement setup for dynamic investigation of the machine tool.</p>
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<p>Schematic representation of the excitation signal during a uniaxial positioning movement at constant feed rate.</p>
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<p>(<b>a</b>) Measured and (<b>b</b>) simulated frequency response functions (FRFs) with linear damping and active motor brake; state S1.</p>
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<p>(<b>a</b>) Measured and (<b>b</b>) simulated FRFs with integrated friction models and active motor brake; state S2.</p>
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<p>(<b>a</b>) Measured and (<b>b</b>) simulated FRFs with active motor brake and with active feed drive control; state S2.</p>
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<p>(<b>a</b>) Measured and (<b>b</b>) simulated FRFs during positioning in X with 100 mm/min feed rate, K<sub>pn1</sub> = 30 Nms/rad, K<sub>pn2</sub> = 15 Nms/rad and K<sub>pn3</sub> = 7.5 Nms/rad; state S2.</p>
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