Strojniški vestnik Journal of Mechanical Engineering Strojniški vestnik - Journal of Mechanical Engineering (SV-JME) Aim and Scope The international journal publishes original and (mini)review articles covering the concepts of materials science, mechanics, kinematics, thermodynamics, energy and environment, mechatronics and robotics, fluid mechanics, tribology, cybernetics, industrial engineering and structural analysis. The journal follows new trends and progress proven practice in the mechanical engineering and also in the closely related sciences as are electrical, civil and process engineering, medicine, microbiology, ecology, agriculture, transport systems, aviation, and others, thus creating a unique forum for interdisciplinary or multidisciplinary dialogue. The international conferences selected papers are welcome for publishing as a special issue of SV-JME with invited co-editor(s). Editor in Chief Vincenc Butala University of Ljubljana, Faculty of Mechanical Engineering, Slovenia Technical Editor Pika Škraba University of Ljubljana, Faculty of Mechanical Engineering, Slovenia Founding Editor Bojan Kraut University of Ljubljana, Faculty of Mechanical Engineering, Slovenia Editorial Office University of Ljubljana, Faculty of Mechanical Engineering SV-JME, Aškerčeva 6, SI-1000 Ljubljana, Slovenia Phone: 386 (0)1 4771 137 Fax: 386 (0)1 2518 567 info@sv-jme.eu, http://www. sv-jme.eu Print: Koštomaj printing office, printed in 275 copies Founders and Publishers University of Ljubljana, Faculty of Mechanical Engineering, Slovenia University of Maribor, Faculty of Mechanical Engineering, Slovenia Association of Mechanical Engineers of Slovenia Chamber of Commerce and Industry of Slovenia, Metal Processing Industry Association President of Publishing Council Mitjan Kalin University of Ljubljana, Faculty of Mechanical Engineering, Slovenia Vice-President of Publishing Council Bojan Dolšak University of Maribor, Faculty of Mechanical Engineering, Slovenia Strojniški vestnik -t!'^ Journal of Mechanical ' Engineering Cover: Mining support steel 31Mn4+QT630. Specimens were cut from a new profile TH-29, cold strained and subsequently heat treated for restauration of original mechanical properties. Figures a) show as-received condition, figures b) after cold straining and heat treatment. Top - microstructures, center - fracture surfaces of tensile test specimens, bottom - Fracture surfaces of Charpy impact test specimens. Image courtesy: University Center for Electron Microscopy, University of Maribor, Faculty of Mechanical Engineering, Maribor, Slovenia ISSN 0039-2480, ISSN 2536-2948 (online) International Editorial Board Kamil Arslan, Karabuk University, Turkey Hafiz Muhammad Ali, King Fahd U. of Petroleum & Minerals, Saudi Arabia Josep M. Bergada, Politechnical University of Catalonia, Spain Anton Bergant, Litostroj Power, Slovenia Miha Boltežar, University of Ljubljana, Slovenia Filippo Cianetti, University of Perugia, Italy Janez Diaci, University of Ljubljana, Slovenia Anselmo Eduardo Diniz, State University of Campinas, Brazil Jožef Duhovnik, University of Ljubljana, Slovenia Igor Emri, University of Ljubljana, Slovenia Imre Felde, Obuda University, Faculty of Informatics, Hungary Janez Grum, University of Ljubljana, Slovenia Imre Horvath, Delft University of Technology, The Netherlands Aleš Hribernik, University of Maribor, Slovenia Soichi Ibaraki, Kyoto University, Department of Micro Eng., Japan Julius Kaplunov, Brunel University, West London, UK Iyas Khader, Fraunhofer Institute for Mechanics of Materials, Germany Jernej Klemenc, University of Ljubljana, Slovenia Milan Kljajin, J.J. Strossmayer University of Osijek, Croatia Peter Krajnik, Chalmers University of Technology, Sweden Janez Kušar, University of Ljubljana, Slovenia Gorazd Lojen, University of Maribor, Slovenia Darko Lovrec, University of Maribor, Slovenia Thomas Lubben, University of Bremen, Germany Jure Marn, University of Maribor, Slovenia George K. Nikas, KADMOS Engineering, UK Tomaž Pepelnjak, University of Ljubljana, Slovenia Vladimir Popovič, University of Belgrade, Serbia Franci Pušavec, University of Ljubljana, Slovenia Mohammad Reza Safaei, Florida International University, USA Marco Sortino, University of Udine, Italy Branko Vasič, University of Belgrade, Serbia Arkady Voloshin, Lehigh University, Bethlehem, USA General information Strojniški vestnik - Journal of Mechanical Engineering is published in 11 issues per year (July and August is a double issue). 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Strojniški vestnik - Journal of Mechanical Engineering 66(2020)12 Contents Contents Strojniški vestnik - Journal of Mechanical Engineering volume 66, (2020), number 12 Ljubljana, December 2020 ISSN 0039-2480 Published monthly Papers Gorazd Lojen, Janez Mayer, Tonica Boncina, Franc Zupanic: Single-Step Heat Treatment for the Restoration of the Mechanical Properties of Cold-Strained Mining Support Steel 31Mn4 687 Wending Li, Guanglin Shi, Chun Zhao, Hongyu Liu, Junyong Fu: RBF Neural Network Sliding Mode Control Method Based on Backstepping for an Electro-hydraulic Actuator 697 Jianyong Liu, Jianfei Sun, Uzair Khaleeq uz Zaman, Wuyi Chen: Influence of Wear and Tool Geometry on the Chatter, Cutting Force, and Surface Integrity of TB6 Titanium Alloy with Solid Carbide Cutters of Different Geometry 709 Xibing Li, Zhe Yu, Xizhao Li, Weixiang Li, Tengyue Zou: Plough-extrusion Forming for Making Micro-groove Heat Pipes on Hydrostatic Thrust Bearings of Heavy Machinery 724 C Satheesh, P Sevvel, R Senthil Kumar: Experimental Identification of Optimized Process Parameters for FSW of AZ91C Mg Alloy Using Quadratic Regression Models 736 Strojniški vestnik - Journal of Mechanical Engineering 66(2020)12, 687-696 © 2020 Journal of Mechanical Engineering. All rights reserved. D0l:10.5545/sv-jme.2020.6818 Original Scientific Paper Received for review: 2020-06-16 Received revised form: 2020-09-09 Accepted for publication: 2020-11-19 Single-Step Heat Treatment for the Restoration of the Mechanical Properties of Cold-Strained Mining Support Steel 31Mn4 Gorazd Lojen1* - Janez Mayer2 - Tonica Boncina1 - Franc Zupanic1 iUniversity of Maribor, Faculty of Mechanical Engineering, Slovenia 2Premogovnik Velenje, d. d., Slovenia The steel 31Mn4+QT630 is used frequently for mining support arches. The supports are cold strained, first during service, and again by re-rolling prior to reinstallation, which results in strain hardening and a loss of ductility. Consequently many of the rerolled arch-segments are not suitable for reinstallation because their mechanical properties are inadequate. The objective of this work was to assess the feasibility of restoration of the required mechanical properties by means of a cost-efficient single step heat treatment. Specimens were cold deformed to different degrees in the range 0 % to 45 % to establish the relation between the degree of cold deformation and the hardness. Differently strain hardened specimens were subjected to subcritical annealing at temperatures 450 °C to 700 °C in the duration 0.5 h to 8 h to determine a suitable time-temperature combinations. Microstructures and mechanical properties were investigated of as-received, cold strained and recrystallized specimens. Tests performed were optical microscopy scanning electron microscopy tensile tests, hardness measurements and Charpy impact tests. Elongation at break of the as-received material was below the requirements of the applicable standard, and its microstructure contained significant fractions of pre-eutectoid ferrite and pearlite. Upon cold straining, hardness increased by approximately 2 HV per 1 % of strain. After one-hour recrystallization at 600 °C to 620 °C, the microstructure consisted of a ferritic matrix containing evenly dispersed globular carbide particles. The original ductility was restored, while the elongation, yield strength, and hardness were higher than in the as-received condition. These results confirmed that it is feasible to restore the original mechanical properties of heavily strained profiles with an adequate single-step heat treatment. Furthermore, they indicated that it should be possible to recover the required properties of inhomogeneously strained material with the same set of well optimized heat treatment parameters. Consequently, it should be possible to increase the reuse rate and to decrease the costs for new support arches significantly. Keywords: microstructure, strain hardening, mechanical properties, recrystallization, steel 31Mn4, steel 1.0520, steel mining supports Highlights • The possibilities were assessed for restoration of the required mechanical properties of damaged (cold strained) mining support steel 31Mn4+QT630. • The relation between hardness and cold deformation in the range from 0 % to 45 % of cold deformation was established experimentally to be approximately linear, 2 HV per 1 % of cold deformation. • Restoration of the required combination of yield strength, tensile strength, elongation, hardness and impact toughness was possible by an adequate single step heat treatment (solely by recrystallization). The optimum combination of time and temperature was established to be about 1 h at 600 °C to 620 °C. • Cold plastic deformation and subsequent recrystallization improved the homogeneity and ductility of the steel compared to new profiles in as-delivered condition QT630. • By the implementation of a simple and inexpensive treatment, consisting of determination of the degree of cold strain and a corresponding single-step heat treatment, it should be possible to increase the reuse rate of mining support arch segments, and, thereby, to decrease the expenses and the environmental impact of mining. 0 INTRODUCTION In underground mines, dynamically capable ground support systems are one of the most reliable options [1]. A steel arch yielding support of roadways is a widely used support structure in European mines [2] and worldwide. For example, in German coal mines, 65 % of roadways are supported by yieldable steel arches and the percentage is still growing [3]. Not only the design of the arch [4], but also the shape of the profiles' cross-sections influence the load capacity of the support. With respect to bearing capacity, the arches built of profiles with open cross-sections like U and V-shaped cross-sections are inferior to those made of concrete filled tubes [5]. However, the significant advantage of yielding steel-arch supports made of V-shaped profiles is the yielding capacity of connections between the segments. The most visible manifestation of strata pressure is the vertical convergence of roadways [6]. When the ground load on the support increases, the sliding of the arch segments over each other facilitates the arch closure, i.e. diminishing the tunnel's radius. This leads to a decrease in the pressure exerted on the arch by the ground, reaching an equilibrium state, overload is prevented and the arch continues to provide support despite significant roadway deformations [7]. Gradual *Corr. Author's Address: University of Maribor, Faculty of Mechanical Engineering, Smetanova ulica 17, 2000 Maribor, Slovenia, gorazd.lojen@um.si 687 Strajniski vestnik - Journal of Mechanical Engineering 66(2020)12, 687-696 stress relaxation (i.e. smooth yielding) decreases the risk of rock burst [8]. To withstand the sufficient clamping forces and to enable smooth yielding, the material must exhibit sufficient strength and hardness. Because the yielding is inevitably accompanied by plastic deformation, adequate ductility is also required to ensure a sufficient yielding capacity. Excessively deformed profiles must be replaced, after the yielding capacity of the connections is fully exploited, or in the case of sudden, heavy overloads. To maintain the production in the underground mines, a huge amount of steel supports must be installed every year. Thus, the new steel supports not only represent a considerable expense, but also have a great impact on the environment. For example, a life-cycle assessment for coal mining operations in Poland revealed that the production of steel supports for underground mining ranks fourth among coal mining related factors that affect greenhouse gas emissions [9]. Therefore, reducing steel consumption by repairing and reusing damaged steel arches not only lowers the costs, but also improves the environmental balance of mining significantly. Currently, about 3*106 kg of steel arch segments are discarded every year in the Coal mine in Velenje. The responsible Department estimates that approximately 1/3 of them could be reused if a reliable and inexpensive heat treatment would be available. The uninstalled supports are divided into two groups. The kinked segments are discarded, while the more uniformly deformed arch-segments can be repaired (i.e. rerolled). However, repair makes sense only if, after re-rolling, the ductility is not too low. Namely, upon re-rolling the total degree of plastic deformation approximately doubles, whereby the ductility decreases. Conveniently, the degree of plastic deformation of rerolled segments is quite uniform along their length. Therefore, the restoration of required mechanical properties should be possible solely by adequate recrystallization annealing, which is much cheaper and less time consuming than new austenitization, quench hardening and tempering. Nevertheless, no reports on attempts to repair of the profiles and no reports on heat treatments of quenched and tempered and subsequently cold strained steels could be found in the literature. Many research groups investigated recrystallization, subcritical, intercritical and/or supercritical annealing of other steels. Zhao and Chen [10] researched a two-stage heat treatment (intercritical and subsequent subcritical annealing) of cold rolled low alloy steel 20CrMnTi in order to obtain fine pearlite and ferrite grains. Optimization of annealing parameters resulted in finer grains, a more uniform distribution of ferrite and pearlite and in an increase of plasticity. Al-Qawabah et al. [11] researched the effect of normalization temperatures in the range 820 °C to 940 °C on the microstructure, microhardness, mechanical behavior and impact toughness of grade 45 low carbon steel. With increasing austenitization temperature, the grain size and the plasticity increased constantly, while the impact toughness and microhardness increased only up to 900 °C and 860 °C, respectively. Kubendran Amos et al. [12] studied the spheroidization of pearlite during subcritical annealing. The effective spheroidization mechanism depended on the initial aspect ratio of a cementite plate. Ji et al. [13] investigated the influences of different temperatures of subcritical annealing on the spheroidization ratio and mechanical properties of medium carbon steel SCM435. With increasing the annealing temperature up to 720 °C, the spheroidization ratio of carbides and ductility increased, while the tensile strength decreased. After annealing in the temperature range 680 °C to 720 °C, mechanical properties were comparable to those obtained with intercritical annealing. Hernandez-Silva et al. [14] studied the spheroidization of cementite in AISI 1541 carbon steel by means of subcritical and intercritical annealing. Spheroidization was faster if prior to subcritical annealing, the steel was cold deformed or intercritically annealed. Bhattacharya et al. [15] studied subcritical annealing of low-carbon microalloyed steel after different sequences of hot rolling. An ultrafine grain microstructure, obtained by 80 % warm deformation followed by intercritical annealing at 800 °C for 5 min and water quenching, resulted in the optimum combination of strength and ductility. It was not reporterd how the optimum combination was determined. Mugas et al. [16] investigated the effects of subcritical annealing followed by oil and water quenching on Ck55 carbon steel. They found out that conventional hardening consisting of austenitization, quenching and tempering, could be replaced by subcritical annealing followed by quenching. Perez [17] explored the impact of different types of annealing treatments on the softening and subsequent work hardening behavior of a martensitic stainless steel. Subcritical annealing resulted in significant softening, while the ductility and work-hardening behavior were not strongly affected. Austenitization followed by slow cooling had no significant influence on hardness and strength. The influences of heating rate during subcritical annealing (recrystallization) of severely deformed 688 Lojen, G. - Mayer, J. - Boncina, T. - Zupanic, F. Strajniski vestnik - Journal of Mechanical Engineering 66(2020)12, 687-696 low carbon steel were explored by Ghiabakloo and Kazeminezhad [18]. Fully recrystallized samples, heated rapidly to 600 °C in a metallic bath, exhibited higher hardness and strength than conventionally recrystallized samples. Yang and Lu [19] reported optimization of intercritical annealing parameters for cold-drawn SCM435 steel wires in order to obtain the optimum mechanical properties of the wire. The degree of cold deformation was not reported, while the best obtainable combination of mechanical properties was found to be the ensile strength 527.1 MPa, hardness 76.2 HRB and ductility 0.437. Yang and Lu [20] and [21] also studied the effects of the type of the protective atmosphere on subcritical and intercritical annealing of SCM435 alloy steel wires. For the improvement of plasticity, a hydrogen atmosphere was found to be more suitable than nitrogen, and intercritical annealing more efficient than subcritical annealing. Joo et al. [22] studied the influence of the shape of the matrix for cold drawing prior to annealing at 705 °C on ductility and spheroidization of carbides in a medium carbon steel for production of wires. Drawing through a non-circular matrix influenced the ductility and spheroidization of cementite beneficially. Influences of the presence of Mn, V and Ti on kinetics and the interaction between recrystallization and precipitation during subcritical annealing of cold-rolled low-carbon steels were investigated by Kapoor et al. [23]. Recrystallization progressed faster in samples containing less Mn, V and Ti. Min and Ha [24] studied the spheroidization of pearlite and mechanical properties of pearlitic steel after cold rolling for 20 % to 40 % and subsequent subcritical annealing at 600 °C and 720 °C for 0.1 h to 32 h. Ductility decreased with increasing cold reduction. Elongation increased dramatically with prolongation of spheroidization annealing. Yang and Liu [25] attempted optimization of the spheroidization of pearlite at 695 °C to 705 °C in AISI 1022 low carbon steel after cold deformation, in order to improve the cold forging properties The temperature and duration of annealing were recognized as the most significant parameters. The optimum values were 705 °C and 8 h. U-shaped profiles and steel 31Mn4 according to DIN 21530-3:2016-09 [26] are used widely for yielding steel arch supports in numerous underground mines all over the world, and the use of 31Mn4 steel is still increasing [27]. However, only a few publications could be found that reported on U-shaped profiles for steel arches, and/or on 31Mn4 steel. Plesea and Radu [28] studied cold bending of U-shaped mining support profiles, residual stress and the occurrence of cracks by means of analytical and numerical analysis. They proved that, already during manufacturing of arch segments by cold bending, very high residual stress can be introduced and cracks can even occur. However, the profiles were not made of 31Mn4 steel, profiles were not quenched and tempered (QT) before bending and the research was focused on production of new arch segments. Radu et al. [29] and Majcherczyk and Malkowski [27] studied the mechanical performance of different support systems and different u-shaped profiles for steel arch supports. Work hardening and heat treatments were not addressed. Kilerci and Culha [30] studied hot forging of 31Mn4. With an increasing number of forging stages, occurrence of cracks and forging forces decreased. Aritan and Can [31] focused on corrosion of 31Mn4 steel arch supports. They compared corrosion rates in underground mine water from a certain mine, in bore water, and in pure water. The corrosion rate was the highest in the mine water. Janas et al. [32] analyzed the load-carrying capacity of steel arch supports made of 31Mn4, and developed software for calculation of load-bearing capacity of profiles TH 29 and TH 34 made of 31Mn4 steel. No data regarding the relationship between hardness and deformation or the properties after recrystallization treatment of strain hardened 31Mn4 steel can be found in open literature neither for normalized nor for quenched and tempered condition. Therefore, the goal of our research was to find a simple, fast, reliable and inexpensive way to determine the actual condition of cold strained segments and to recover the ductility of excessively strain hardened segments. 1 METHODS Cuboid samples were machined from new (unused) steel arch segments. The samples were compressed to different degrees, and their hardness was measured in order to determine the relationship between the hardness and the degree of cold deformation. Subsequently, the samples were recrystallized using different time-temperature combinations to establish a suitable combination for the restoration of original mechanical properties. Tensile tests and Charpy impact tests were performed to verify the suitability of the heat treatment. Evolution of microstructure upon strain hardening and subsequent heat treatment and fracture surfaces were investigated with a high resolution scanning electron microscope. Single-Step Heat Treatment for the Restoration of the Mechanical Properties of Cold-Strained Mining Support Steel 31Mn4 689 Strajniski vestnik - Journal of Mechanical Engineering 66(2020)12, 687-696 2 EXPERIMENTAL Samples were obtained from new profiles TH 29 made of steel 31Mn4, delivered in quenched and tempered condition +QT630, which is included in the standard DIN 21530-3:2016-09. The chemical composition of the analyzed profile was found to be within the specified range, Table 1. An ARL 3460 optical emission spectrometer was used for the chemical analysis. Table 1. Chemical composition of steel 31Mn4 to DIN 215303:2016-09 and analyzed composition Element DIN 21530-3:2016-09 [wt. %] Analyzed [wt. %] C 0.28 to 0.36 0.36 Si 0.20 to 0.50 0.36 Mn 0.80 to 1.10 0.91 P < 0.035 0.016 S < 0.035 0.007 Al < 0.020 - Cu < 0.35 - Ni 0.02 Cr 0.05 Metallographic examinations, hardness tests, tensile tests and Charpy impact toughness tests followed. Examinations and tests were done with specimens in different conditions: As-received (+QT630), strain hardened by compression, and in recrystallized condition. Classic metallographic preparation was applied, consisting of grinding, polishing and etching with the nital/picral, ratio 10:1. The microstructures were examined with a scanning electron microscope FEI Sirion 400 NC, equipped with the EDX system Oxford INCA 350. Vickers hardness HV 30 was measured with a VPM Leipzig HPO 250 hardness tester. The shaded area in the Fig. 1 indicates the position of specimens in the TH 29 profile. The longitudinal axis of specimens was parallel to the profiles' longitudinal axis. Specimens 15 mm x 15 mm x 30 mm were used to determine the relation between the degree of cold plastic deformation and hardness. The cuboids were compressed in the longitudinal direction with a universal hydraulic testing machine VPM Leipzig EDZ 100. Preliminary tests confirmed that the highest increase of hardness occurred in the area around the body center of a compressed specimen and that the results are practically independent of the section-orientation. Therefore, the hardness was measured in this area. Afterwards, the same specimens were used to assert the appropriate time and temperature of recrystallization. 150.5 mm± 2.5 Fig. 1. Cross-section of the profile TH 29. The shaded surface at the bottom shows the position of specimens Cuboids 15 mm x 30 mm x 90 mm and 15 mm x 30 mm x 60 mm were used for the tensile tests and Charpy impact toughness tests, respectively. Some were left in as-received condition, while others were compressed in the transverse direction with a hydraulic compression testing machine VPM Leipzig ZDM 300. Some compressed specimens remained in the cold strained condition, while the rest were recrystallized. Adequate test pieces were machined from specimens in all three conditions: Short standard B-type specimens according to DIN 50125 [33] for tensile tests, and ISO-V-notch specimens according to EN ISO 148-1 [34] Charpy impact toughness tests. A Messphysik Beta 300 machine was used for tensile tests, and a 300 J notch impact tester PSW 300 from VPM Leipzig for the Charpy tests. The Vickers hardness of V-notch specimens was measured additionally. 3 RESULTS AND DISCUSSION 3.1 Relationship between Deformation and Hardness The 15 mm x 15 mm x 30 mm cuboids were compressed in the longitudinal direction to different degrees of final deformation to investigate the relationship between the degree of cold deformation and hardness. Hardness was measured in the area of the highest deformation, on the cross-section through the body center. The results are presented in Fig. 2. The relationship between the degree of deformation and hardness was almost linear. Assuming a linear relationship, shown by the calculated dotted trend line, the hardness increased by 2.03 Vickers units per 1 % cold deformation, Eq. (1): HV = HV0 + 2.03 x e . (1) 690 Lojen, G. - Mayer, J. - Boncina, T. - Zupanic, F. Strajniski vestnik - Journal of Mechanical Engineering 66(2020)12, 687-696 In the Eq. (1), HV represents the final hardness after cold deformation, HV0 the initial hardness before cold deformation, and s the degree of cold deformation in percentage. 340 320 300 o r-j X 280 f | 260 X 240 220 0 10 20 30 40 50 Degree of cold strain [%] Fig. 2. Influence of cold straining on the hardness of 31Mn4+QT630 steel. The exact degrees of compressive cold deformation were 0 %, 7.8 %, 17.6 %, 30.8 % and 43.9 % Opposite to our specimens, steel arches in the mines are deformed by bending. Therefore, the maximum plastic deformation always occurs on the surface of the profile, where the hardness can be measured easily and quickly with portable hardness testers. Hence, Eq. (1) can be used for determining the degree of cold plastic deformation of profiles. 3.1 Recrystallization Parameters In the first step, two groups of specimens, cold strained for 10 % and 45 %, were annealed for 1 h at different temperatures. The specimens were inserted into a preheated furnace and time was measured from the moment when the specimens reached the set temperature. The necessary time for heating up was determined experimentally. The test pieces were used with a drill hole reaching into the body center. One thermocouple was placed into the drill hole, and the other was in contact with the surface. For both thermocouples, the necessary time to reach the pre-set furnace temperature were measured; t1 for the surface and t2 for the body center. The delay of the body center At was determined from their difference. Because the test pieces with a drill hole were inconvenient for other experiments, later, only the surface temperature was measured. After t1 + A t, we started to measure the soaking time. It was found that a hardness of about 250 HV can be obtained at approximately 600 °C, Fig. 3a. The less strained specimens softened more slowly. Hence, the risk that, after recrystallization, the less strained sections of profiles would be too soft, while more heavily strained areas would regain the required mechanical properties, is low. In the second step, a third group of specimens, all of them being cold strained for 45 %, was annealed at various temperatures for up to 8 hours to assess the influence of the annealing time, Fig. 3b. Again, the optimum temperature was found to be around 600 °C. At 600 °C, the hardness reached 259 HV 30 in 0.5 h, 254 HV 30 in 1 hour and 242 HV 30 in 2 hours. It was evident that temperatures below 600 °C were too low to obtain hardness near 250 HV 30 in a reasonable time. On the other hand, temperatures of 650 °C and above were too high. At 650 °C, the hardness reached 350 330 310 290 270 250 230 210 190 170 150 I 45 % 10% a) 400 500 600 700 Temperature [nC] 350 330 310 290 £ 270 I 250 I 230 210 X 190 170 150 b) k450°C IWson °c >50 °C 600 °C P^so^c i 700 °C 0 1 2 3 4 5 6 7 8 Annealing time [h] Fig. 3. Hardness of 31Mn4+QT630; a) cold strained for 10 % and 45 % and annealed for 1 h at different temperatures, and b) 45 % cold strained and annealed at different temperatures for up to 8 hours Single-Step Heat Treatment for the Restoration of the Mechanical Properties of Cold-Strained Mining Support Steel 31Mn4 691 Strajniski vestnik - Journal of Mechanical Engineering 66(2020)12, 687-696 250 HV 30 in about 15 minutes. Such short times are unsuitable for the industrial environment, because the process is too difficult to control with sufficient accuracy. 3.3 Mechanical Properties The results of mechanical tests are summarized in Table 2. The minimum mechanical properties for 31Mn4+QT630 steel according to DIN 21530-3:201609 are listed at the bottom of the Table. 3.3.1 Tensile Properties The tensile tests revealed that tensile strength and yield strength of as-received specimens met the requirements of DIN 21530-3, while elongation of one specimen did not reach the required 16 %. Like the low hardness, the insufficient elongations can also be attributed to the inhomogeneous microstructure, which is discussed in Section 3.4.1 Microstructures. As expected, tensile strength increased strongly due to strain hardening, while elongation at break decreased. All recrystallized specimens had yield points far above the requirements of the Standard. One of them, the one strained for 33 % and recrystallized at 620 °C, exhibited insufficient tensile strength. The tensile strengths of the recrystallized specimens were lower than in the as-received condition, while their yield points were, on average, significantly higher. The elongations at break of the recrystallized specimens were, on average, also higher than in the as-received condition. Only one of them, strained for 45 % and recrystallized at 600 °C, did not quite reach the required 16 % elongation. These test results imply that it is not easy to find the right balance between tensile strength and elongation. Moreover, the recrystallization temperature should be chosen carefully with respect to the degree of previous cold plastic deformation. In addition, it was observed that the characteristics of the yield point changed upon recrystallization. Non-strained specimens and strained but not recrystallized specimens exhibited an offset yield point Rp02, typical for hardened and cold worked steels. Recrystallized specimens exhibited a higher yield point Reh and lower yield point Rel, which is characteristic for mild steels. The appearance of Reh and Rel on the one hand, and Reh being higher than Rm on the other hand, can be attributed to the diffusion and piling of foreign atoms around the dislocations in the iron lattice [35] during the recrystallization treatment, which hindered the start of the dislocation glide during the tensile test. 3.3.2 Hardness The hardness of the ISO-V and the tensile test specimens in as-received condition was not measured, because the hardness in this condition was already determined with other specimens, and was about 240 HV 30 on average, Table 2. After recrystallization, the average hardness was higher than in the as-received Table 2. Mechanical properties of 31Mn4+QT630 steel in as-received, cold strained, and recrystallized condition, and minimum requirements according to DIN 21530-3:2016-09 s [%] T [°C] tr [h] Reh [MPa] Rel [MPa] Rp02 [MPa] Rm [MPa] A5 [%] Av [J] HV 30 0 - - - - 727 868 16.9 100a) 241 13 - - - - - - - 99a) 296 30 - - - - 1067 1180 6.8 42a) 296 45 - - - - 1132 1202 5.3 40a) 339 13 600 1 - - - - - 117a) 246 28 600 1 829 800 - 836 19.9 - - 29 600 1 867 804 - 849 16.3 - - 30 600 1 878 793 - 834 17.7 91a) 266 45 600 1 859 830 - 851 14.6 99a) 285 31 620 1 844 787 - 818 20.9 - - 33 620 1 766 699 - 715 16.4 - - DIN 21530, QT630 > 630 - - > 790 > 16 >70/75b) - s = cold strain [%]; Tr = recrystallization temperature [°C]; tr = recrystallization time [h]; Reh = upper yield point; R>i = lower yield point, Rp02 = offset yield point (at 0.2 % of plastic deformation); Rm = Tensile strength, A5 = elongation after fracture, obtained with test pieces having a length:diameter ratio 5:1; Av = impact energy absorbed in the Charpy test; HV30 = Vickers hardness, measured with a load of 30 kg; a) ISO-V specimens at room temperature; b) DVM specimens, single specimen/average of 3 specimens at room temperature 692 Lojen, G. - Mayer, J. - Boncina, T. - Zupanic, F. Strajniski vestnik - Journal of Mechanical Engineering 66(2020)12, 687-696 condition. Only one specimen, strained for 13 % and recrystallized at 600 °C, exhibited a hardness lower than 250 HV. Nevertheless, even this specimen was harder than the majority of specimens in the as-received condition. 3.3.3 Impact Toughness For the average of three specimens, the standard DIN 21530-3:2016-09 specifies the threshold value of 75 J DVM. However, ISO-V specimens are more common than DVM specimens. Therefore, the ISO-V form was selected for impact toughness tests in our research. A comparison of DVM and ISO-V values [36] shows that 63 J DVM equals 63 J ISO-V. Above 63 J, the ISO-V-values are higher than DVM-values: 72 J DVM corresponds to 75 J ISO-V, and 77 J DVM corresponds to 82 J ISO-V. Accordingly, the threshold value for ISO-V specimens would be around 80 J. In this regard, the impact toughness of recrystallized specimens was always above the required minimum value, and even a moderately cold strained specimen (13 %) exhibited impact toughness higher than required. As expected, the impact toughness of heavily strained (30 % and 45 %) and not recrystallized specimens was insufficient. 3.4 Microstructures and Fracture Surfaces Microstructures and fracture surfaces of specimens were examined in all three conditions. However, for the purposes of this examination, the most relevant comparisons are between the as-received and the recrystallized conditions. Therefore, only photographs of these two conditions are presented below, while the cold strained condition is only described shortly in the text. 3.4.1 Microstructures Characteristic microstructures of as-received, cold deformed and recrystallized material are presented in Fig. 4. The SEM image of the as-received material, Fig. 4a, reveals that the microstructure contained ferritic grains (dark), relatively coarse lamellar areas, and areas containing much finer lamellar and globular carbides. The coarser lamellar areas were pearlite. The presence of proeutectoid ferritic grains and pearlite indicates that the steel was not quenched properly - the cooling rate was too low. The presence of ferrite explains the low hardness of the as-received material, and the inhomogeneity of the microstructure contributed to the relatively low elongations. Because the material was tempered after quenching, the areas containing finer carbides could be bainite and/or tempered martensite, as well as partially spheroidized fine pearlite. Fig. 4. SEM SE micrographs; a) as-received, and b) 45 % cold strained and recrystallized 1 h at 600 °C After cold straining, the appearance of carbide-rich areas did not change much. As expected, the majority of deformation took place in the ferritic regions, because the yield stress of ferrite is lower than that of carbide-containing areas. Due to recrystallization, the homogeneity of the material improved significantly. The shape of cementite platelets transformed from lamellar to nodular, while the already more or less globular particles coarsened slightly, Fig. 4b. However, some areas could still be observed containing solely very fine cementite particles. These areas are former martensite and/ or bainite, while areas containing predominantly coarser particles are most likely spheroidized pearlite. A microstructure consisting of a ferritic matrix containing fine dispersed nodular cementite particles is normally representative for the quench-hardened and tempered steels. However, as shown in Fig. 4b, cold plastic straining followed by recrystallization can also produce uniformly dispersed nodular cementite. Single-Step Heat Treatment for the Restoration of the Mechanical Properties of Cold-Strained Mining Support Steel 31Mn4 693 Strajniski vestnik - Journal of Mechanical Engineering 66(2020)12, 687-696 3.4.2 Fracture Surfaces The fracture surfaces of tensile test specimens are presented in Fig. 5. It is evident that the fractures of the as-received and recrystallized specimens, Figs. 5a and b, respectively, were ductile. The fracture surfaces in both conditions exhibited small dimples of fairly uniform size, though, the surface of the as-received specimen was slightly less uniform, which is, as the lower elongation, consistent with its less homogeneous microstructure. The ductile fracture still prevailed even in the cold strained condition. However, the dimples were larger, much less uniform and some brittle fractured areas were observed. Such fracture behavior resulted in significantly lower elongations. Fig. 5. The fracture surfaces of the tensile test specimens, SEM SE images; a) as-received, and b) 30 % cold strained and recrystallized 1 h at 600 °C Compared to the tensile test specimens, the fracture surfaces of the Charpy impact test, Fig. 6, specimens were less uniform, and all of them showed the mixed fracture character. In the as-received specimen, Fig. 6a, two large cavities were observed still containing spherical particles - manganese sulfide inclusions, as the EDX analysis revealed. Similar cavities were also present in the cold strained and recrystallized specimens, Fig, 6b. The surfaces of these cavities were smooth, as is characteristic for a transcrystalline brittle fracture. The areas around the cavities were rougher, showing a more ductile behavior. Also, outside the large cavities, flat surfaces could be observed, the least of them in the recrystallized specimen. This indicates that the rupture started with a brittle fracture by the instant formation of relatively large cracks around the sulfide inclusions and continued in a partially ductile way. Fig. 6. Fracture surfaces of Charpy impact test specimens, SEM SE micrographs; a) as-received, and b) 45 % cold strained and recrystallized 1 h at 600 °C Comparing the Figs. 6a and b, the proportion of brittle fracture seems to be slightly larger in the as-received condition than after recrystallization. The above observations are consistent with differences in microstructures. However, the absorbed impact energies of the as-received and recrystallized specimens were very similar. This indicates that the differences in their microstructures did not influence their impact toughness crucially - most likely because of the presence of numerous inclusions. 694 Lojen, G. - Mayer, J. - Boncina, T. - Zupanic, F. Strajniski vestnik - Journal of Mechanical Engineering 66(2020)12, 687-696 4 CONCLUSIONS Microstructures and mechanical properties of as-received, cold strained and recrystallized specimens of 31Mn4+QT630 steel, cut from a TH 29-profile, were investigated to assess the possibilities of treatment and reuse of heavily strained segments of a steel arch support system. The results can be recapitulated as follows: • Apart from elongations, the mechanical properties of as-received material corresponded to the applicable standard. Elongations were partially below the requirements for the condition +QT630. The hardness was below 250 HV 30. • The microstructure of the as-received material was not characteristic for a QT steel. It was inhomogeneous, and contained significant fractions of proeutectoid ferrite and pearlite. • Upon cold plastic straining, hardness increased approximately linearly with the degree of strain, approximately 2 HV per 1 % of strain. Thereby, if the original hardness is known, simple and reliable estimation of the degree of cold plastic strain is possible. • The homogeneity of the material was improved significantly upon cold plastic straining and subsequent recrystallization. The microstructure of recrystallized specimens consisted of a ferritic matrix containing evenly dispersed globular carbide particles, which was similar to microstructures, characteristic for properly quenched and tempered carbon steels. • By recrystallization of cold strained specimens for 1 h at 600 °C to 620 °C, the required combination of mechanical properties could be restored in most cases. On average, yield strength, elongation and hardness were higher and the tensile strength was lower than in the as-received condition. Impact toughness reached values over 90 J ISO-V, and was on average practically the same as in the as-received condition. • By means of an adequate single step heat treatment, the initial mechanical properties of strain-hardened steel can be recovered or even improved. • By the implementation of a treatment protocol, consisting of determination of the degree of cold strain and a corresponding single-step heat treatment, it should be possible to increase the reuse rate of steel arch segments and thereby to decrease the expenses for new profiles significantly. 6 ACKNOWLEDGEMENTS This work was supported by the coal mining company Premogovnik Velenje d.d. The sponsor had no influence on the study design, data collection, interpretation of results, and decision to submit the article for publication. 7 REFERENCES [1] Potvin, Y. (2009). Strategies and tactics to control seismic risks in mines. The Journal of The Southern African Institute of Mining and Metallurgy, vol. 109, p. 177-186.. 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D0l:10.5545/sv-jme.2020.6866 Original Scientific Paper Received for review: 2020-07-19 Received revised form: 2020-10-19 Accepted for publication: 2020-11-06 RBF Neural Network Sliding Mode Control Method Based on Backstepping for an Electro-hydraulic Actuator Wending Li123* - Guanglin Shi1 - Chun Zhao23 - Hongyu Liu23 - Junyong Fu23 Shanghai Jiao Tong University, School of Mechanical Engineering, China 2Shanghai Aerospace Control Technology Institute, China 3Shanghai Servo-system Engineering Research Center, China Aiming at the interference problem and the difficulty of model parameter determination caused by the nonlinearity of the valve-controlled hydraulic cylinder position servo system, this study proposes a radial basis function (RBF) neural network sliding mode control strategy based on a backstepping strategy for the electro-hydraulic actuator. First, the non-linear system model of the third-order position electro-hydraulic control servo system is established on the basis of the principle analysis. Second, the model function RBF adaptive law and backstepping control law are designed according to Lyapunov's stability theorem to solve the problem of external load disturbance and modelling uncertainty, combined with sliding mode control strategy and virtual control law. Finally, simulation and experiment on MATLAB Simulink and semi-physical experimental platform are accomplished to show the effectiveness of the proposed method. Moreover, results show that the designed controller has high tracking accuracy to the given signal. Keywords: RBF neural network, sliding mode, backstepping, non-linear control, electro-hydraulic actuator Highlights • The third-order non-linear model of electro-hydraulic servo system is established with disturbance. • The observer and the sliding mode term are used to remove the different disturbances. • By introducing the backstepping method, the virtual control quantity is solved first, and then the control law is solved by the neural network sliding mode algorithm, which achieves high control precision. • The simulation analysis of high-speed and low-speed operating conditions was carried out, and the experimental results were compared with the simulation results, which showed good consistency. 0 INTRODUCTION The electro-hydraulic servo control system is widely used in aerospace, national defence, civil industry, robot [1], and other fields because of its large power-weight ratio, rapid response speed, strong bearing capacity, and high control accuracy [2]. Electro-hydraulic servo actuators are often used as the execution unit, such as the thrust vector control of rockets and missiles, attitude control systems of aircraft and tanks, stability control systems of vehicle active suspension, and control systems of mill presses [3]. As a key executive system, the performance of the electro-hydraulic servo actuator directly influences the precision, stability, and reliability of the control system [4]. The requirements of the electro-hydraulic control system are also continually improved in terms of mechanical working accuracy [5], response speed, and automation degree [6]. This improvement requires not only the high performance of hydraulic control components but also involves the use of advanced control strategies. However, the high-performance requirement of electro-hydraulic servo system (e.g., complexity, nonlinearity, parameter uncertainty, and load variability) cannot be achieved by using the traditional control methods. The use of intelligent control methods in solving the control problem of the electro-hydraulic servo system is a key technical issue in the development of high-performance electro-hydraulic servo systems [7]. In recent years, many scholars have applied advanced control methods, such as adaptive, robust, sliding mode, and intelligent controls [8], to the electro-hydraulic servo control system to improve the control performance of the system, and have achieved varied results [9]. However, the work environment of the electro-hydraulic servo system of a launch vehicle is complex, that is, the load characteristics have many new characteristics, and the stability of the rocket control system requires the launch vehicle to have good dynamic characteristics [10]. Therefore, other effective control strategies need to be explored [11]. The backstepping method, which refers to a step-by-step method of designing control laws, is effective in solving non-linear control problems [12]. Moreover, many scholars have investigated and applied this method in non-linear system control [13]. Neural network control is a new strategy in overcoming the control problem of systems that are complex, nonlinear, or uncertain [14]. Sliding mode control is a *Corr. Author's Address: School of Mechanical Engineering, Shanghai Jiao Tong University, 800 Dongchuan Road, Minhang District, Shanghai 200240, China, liwending1983@163.com 697 Strajniski vestnik - Journal of Mechanical Engineering 66(2020)12, 697-708 special kind of non-linear control that designs the switching hyperplane of the system according to the expected dynamic characteristics of the system [15]. In the present study, a three-order sliding mode control of the neural network method of the radial basis function (RBF) is proposed on the basis of the backstepping strategy for the servo system of valve-controlled electro-hydraulic position to solve further the non-linear modelling uncertainty and interference of the combined neural network, sliding mode, and backstepping technologies [16]. 1 NON-LINEAR MATHEMATICAL MODEL AND PROBLEM DESCRIPTION In this study, the valve-controlled electro-hydraulic servo system is mainly composed of a controller, a motor pump power source, an accumulator, a pressure sensor, a relief valve, a servo valve, a hydraulic lock, a position sensor, and a hydraulic cylinder [17], as shown in Fig. 1. Fig. 1. The electro-hydraulic actuator system The system block diagram of the electro-hydraulic servo system is shown in Fig. 2. Fig. 2. Block diagram of the electro-hydraulic actuator In Fig. 2, r(t) is the reference input or set value of the system, e is the control deviation, u is the control law, y is the system output, d(t) is the continuous external disturbance, and k is the feedback coefficient. By measuring the feedback signal in the position, a closed loop of the negative feedback deviation control is formed to meet the requirements of the output signal. The kinematic equation of the system is determined by considering the non-linear factors. my = Pla - By - AfSf(y) + f(y y > 0 (1) where m is equivalent to the total weight of the load on the piston rod, y is the piston rod output displacement, pL is the pressure difference between two cavities of the actuator, A is the area of action of the liquid pressure, B is the viscous damping coefficient, Af is the Coulomb friction amplitude, Sf (y) is the Coulomb friction shape function, and f (y, y, t) is the unmodeled dynamic and external disturbance function. The pressure dynamic equations of the actuator are expressed as follows [18]: ß Pi = Voi + Ay P 2 = V02 - Ay (- Ay - C,pL + Q), (Ay + C,pL - 02), (2) (3) Qi =4lkqXxv[0(Xv ps - Pi +©(-xv Pi - Pr ], (4) where p1 and p2 are the pressure values of the two cavities of the actuator, V01 and V02 are the initial volumes of the two cavities of the actuator, ft is the bulk modulus of elasticity of oil, Ct is the leakage coefficient, and Q1 and Q2 are the flow rates of the two cavities of the actuator. The flow equation of the servo valve to the two-cavity actuator is presented as follows: Q =42kqX xv [0(Xv )4 Ps - PX +®(-xv )V Pi - Pr ], (4) Q2 =4lkql xv [&(xv y p2 - pr +®(~xv V Ps - p2 ], (5) where kq1 and kq2 are the servo valve flow gains, xv is the servo valve spool opening, ps is the oil inlet pressure, and pr is the oil outlet pressure. The value of 0() can be expressed as follows: ©(•) = Assume that kn=kn- 1, ■ > 0 0, <0' (6) .lq kq1 = kq2. The dynamic servo valve can be simplified to a proportional link because the frequency width of the servo valve is remarkably higher than that of the system, xv=k-u. In this case, 0(xv) = 0(u). 698 Li, W.D. - Shi, G.L. - Zhao, C. - Hongyu Liu, H.Y. - Fu, J.Y. Strajniski vestnik - Journal of Mechanical Engineering 66(2020)12, 697-708 If ku = J2kqki, and Aj = [0(wps - p1 + &(-up1 - pr ] , and A2 = [©(mp2 - pr + ©(-wps - p2 ] then the flow equation can be expressed as: Qj = ku Aj u, Q2 = ku Aj u. By combining Eqs. (2) and (3) we get: P - P2 = (— + —)hpu- (— + —)Apx2 12 V, V2 u V V2 2 -(V+. (7) Set the state variables to x = y x2 = y , and x3 = yp - p2)A =pLA, and the state equation of the electro-hydraulic servo actuator is expressed as follows: x2 = — X3 - — Bx2 - — AfSf (x2) + — f (y, y, t) m m m m Aj A 1 1 X3 = (—- + —)Kßu- (— + — )A ßx. (8) Vj V2 V— V2 -1 + -1 )ßC,pL + d2(t) '1 V2 2 To simplify the expression, let fl(x) = — Bx2 + — AfSf ( x2) 2 STRATEGY SELECTION AND CONTROLLER DESIGN 2.1 Strategy Selection The backstepping method is an effective non-linear control method that emerged in the 1990s [19]. This method can solve the problem of the lack of constructivism in the traditional Lyapunov function [20]. According to the structural characteristics of the controlled object, the backstepping method establishes the Lyapunov function of the entire system via progressive recursion and realizes the stability control of the non-linear closed-loop system [21]. The RBF neural network algorithm can approximate any non-linear network [22], and the sliding mode control algorithm can improve the robustness of the system by selecting switching functions [23]. The RBF neural network is a kind of a single hidden layer of the three-layer feedforward network with input and output non-linear mappings. The network structure is shown in Fig. 3. Fig. 3. RBF neural network structure g(u, x) = + VjL)kuAß , and '1 '2 Ux) = (V + V-)+ V + yWC^L,, and '1 V2 '1 V2 di(t) = fl ( Xl, X2 , t ) Then, Eq. (8) has the following form: x1 = X2 1 X2 = — X- m ' X- = g (u, x)u - f ( x) + d2 (t ) (9) Eq. (9) shows that strong nonlinearity exists, and building an accurate model of the system is difficult. At the same time, the external interference has a substantial influence on the system. h- = exp(-- x - c 2b? / = W*Th f ( x) (10) (11) where fu is the ideal RBF network output, x is the RBF network input, j is the jh node of the network hidden layer, h = [/?j]T is the Gaussian basis function output of the network, W* is the ideal weight of the network, bj is the base width parameter of node j, bj> 0, Cj is the centre vector of the network node j, and sf is the network approximation error. Sliding mode control is a discontinuous kind of non-linear control method for designing a stable motion mode, that is, the sliding mode surface for the system state or error in advance, and creating the controller to guide the system trajectory to this preset mode and force the system trajectory to maintain and move along the mode in the future [24]. The system is defined as x = f(x, u, t) if the plane of state s(x) and 2 171 RBF Neural Network Sliding Mode Control Method Based on Backstepping for an Electro-hydraulic Actuator 699 Strajniski vestnik - Journal of Mechanical Engineering 66(2020)12, 697-708 control function u(x) are available. The sliding mode control can be expressed as follows: lu+ (x), i(x) > 0 + _ u = < u (x) ^ u (x). (12) \u_ (x), i(x) > 0 The system triggered from any initial state can reach the state plane s(x) = 0 within a certain period, and the steady-state point of the system is reached under the action of u(x). 2.2 Controller Design A state observer of disturbance d1(t) is designed according to the method in [25], xe = dj(t). x\= x 2+ 3a> ^ - Xi j X2= — -fi( x) + x,+ 3œ2( Xi- à), m v > xe =©3 ^ x1—xi j (13) dl(t) is the observation error. Eq. (9) can be written as follows: x2 = — x3 - fi(x) + di(t) + dl(t). (14) m x3 = g(u, x)u - f2 (x) + d2 (t) The state error of the system is defined as: e2 = e + ke = x2 - x2d , (15) In order to alleviate the discontinuity of sliding mode switching and solve the problem of indifferentiability of sign function, the sign function sgn(s) is replaced by 2 / narctan(Ks). K is a positive real number, and the greater the value, the more accurate the approximation. When K= 300, the approximate curve of sign function represented by 2 / n arctan(Ks) is shown in Fig. 4. Fig. 4. Approximation curve of the sign function The virtual control rate a can be expressed as: a = -m(celel - f (x) - x2d + di (t) + ¡usx ) 2 --xfy m arctan( Ktst ), K (19) where, Kj is a sufficiently large positive real number. The Lyapunov function is defined as [27]: V = 1 / V 2 (20) sj = -ßs1 - — x ^ arctan(^sj) + dj(t), (21) where, x1d is the input of the system, x2d = x1d- k1e1, and a is the virtual control quantity of x3. The error can be written in this form to avoid differential explosion. 1. Step 1: Find the virtual control law [26] The sliding mode surface is defined as Eq. (16): S1 = Ceiei + e2 . (16) Substituting Eqs. (14) and (15) into Eq. (16): ¿1 = ceiei + — x3 -fi(x) + di(t) + di(t) - xld. (17) m An appropriate positive real number n1 is selected to design the control law based on disturbance observation and sliding mode. The virtual control quantity a is: a = -m(ce1è1 - f1( x) - x2d + di(t ) + sgn(s). (18) Vj = -ßsf--tsl arctan(^jSj) + sldl(t). (22) When s e (-<» , <»), it can be proved that: f I Sl I ~ K ~ Sl arCtan( Kl°l) " f I (23) According to Eqs. (22) and (23) the following equation can be obtained: 2 vi ^ 1 si 1 + sidi(t) nK1 2 2 nK1 (24) According to the method in reference [28], the inequality in Eq. (24) can be solved as follows: 700 Li, W.D. - Shi, G.L. - Zhao, C. - Hongyu Liu, H.Y. - Fu, J.Y. Strajniski vestnik - Journal of Mechanical Engineering 66(2020)12, 697-708 V(t) < n (25) V1(t) converges asymptotically, and the law of convergence is affected by the values of n, K1 and p. 2. Step 2: Find the system control law Definition of sliding surface: (26) S2 deiei de2e2 ^ e3 • The derivative of Eq. (26) is: s2 = delel + de2e2 + g(u,x)u - f2(x) + d2(t) -a. (27) Design control law u: 1 -\~deA - de 2e2 + f2(X) g(u, X) +a arctan(K2s2)]. (28) Due to the existence of nonlinearity, the mathematical models of f2(x) and g(u,x) cannot be accurately obtained and the control law cannot be solved. Therefore, f2(X) and g(u,x) need to be estimated or approximated. RBF neural network learning method is adopted to approximate functions f2(x) and g(u,x), then: fg(u, x) = Wghg (x) + £g [ f2( x) = W^h f (x) +Ef (29) Wg and W n are respectively the dynamic estimated weights of the neural network approximation of function g(u, x) and f2(x). The neural network approximation of g(u,x) and f2(x) is written in the following form: g (u, x) = W g h g (x) (30) [ f 2( x) = W f 2h f ( x) The control law, Eq. (28), is written as follows: 1 u = —-- [~delex - de2e2 + f 2 (x) g (u, x) + a arctan(K2s2)]. (31) The observation errors are: g(u,x), f2(x), g (u, x) = g (u, x) - g (u, x), (32) f2(x) = J 2( x) - f2( x), (33) s2 = delel + de2e2 - g(u, x)u + g(u,x)u - f2(x) + d2(t) -a, (34) where, tx = -m |ce1e1 + f (x) + x2d + d[r\1 arctan(^1s1)]|. Substitute Eq. (31) into Eq. (34): S2 = -[g(U x) - g(u, x)]u + [ J2 (x) - f2(x)]-^2 arctan(^2s2) + d2(t), (35) Define W. = W f2 - W!,, W„ = W. - W*. Eq. (35) can be written as follows: ¿2 = Wf2 hf2 (x) -sf2 -[ Wghg(x) - Sg ]u - -q2 arctan(K2s2) + d2 (t). Take the Lyapunov function as: (36) L 2 ^ ■ — WT Wf + — Wg' Wg. (37) 2Yf2 f2 f2 27g g g ' ' The derivative of Eq. (37) is: ü = Wf2[s2h (x)- — W t2 ] 2 2 Xf2 1 -Wa [S2h a (x)uH W g ] y9 g L 2 g ' +s2[-e h+egu- 772 arctan(K 2 s2) + d 2(t)j. (38) The adaptive law of Wf and Wg is: |W r2 = yh s2h f2( x) [w^=-7gs2h g (x)u (39) Eq. (37) is written as follows: L2 = s2 [-£fi + sgu - arctan(^2s2) + d2 (t)] = S2[S/2 + sgU + d2(t)]-^2 | s2 |. (40) When the approximation error ef and eg is limited to a small enough range, if n2 satisfies n2> |-+ egu+d2(t), then L2 < 0. When L2 = 0 , s2 ^ 0. According to Lasalle's invariant set principle, the system is asymptotically convergent and globally stable. t, s2 ^ 0. Since L2> 0 and L2 < 0, then when t ^ <» , L^ is bounded. 3 SIMULATION AND ANALYSIS A system simulation model is built in Simulink to verify the effectiveness of the proposed control Substitute Eq. (32) into Eq. (27): RBF Neural Network Sliding Mode Control Method Based on Backstepping for an Electro-hydraulic Actuator 701 u Strajniski vestnik - Journal of Mechanical Engineering 66(2020)12, 697-708 strategy in this study [29]. The simulation parameters are listed in Table 1. Table 1. System parameter No. Symbol Description Values 1 A Hydraulic area 0.01 m2 2 Vo Initial volume of cavity 0.00124 m3 3 m Total load 2000 kg 4 B Viscous damping coefficient 95 N-s/m 5 p Oil density 850 kg/m3 To prove the effectiveness of the new control strategy, a proportional-integral-derivative (PID) controller is compared [30]. 1. The PID controller: kp = 1500, k= 10, kd= 0; 2. RBF neural network sliding model control with extended state observer (RBFSMESO): k1 = 3, k2 = 2, , = 5, del = 6, de2 = 1, y& = 12, he 2.5, The simulation analysis of low-speed tracking and weak external interference, high-speed tracking and strong external interference is carried out below. 1. Low-speed tracking and weak external interference: Expected curve of trajectory tracking: x1d = 0.015sin(0.159x 2nt)[1 - exp(-10t3)], (41) and external load interference: fd1 = 0 t < 5 s fd2 = 20000{1 - exp[-0.1x (t - 5)3]} t > 5 s fd3 = 5000sin(2^t) t > 15 s (42) K, = K2 = 1000. The track tracking performance and tracking error of PID and RBFSMESO controllers are shown in Figs. 5 and 6, respectively. It can be seen from the figures that the expected speed is slow (approximately 0.159 Hz), and the external interference is small Fig. 5. PID tracking performance and tracking error Fig. 6. Tracking curve of backstepping method 702 Li, W.D. - Shi, G.L. - Zhao, C. - Hongyu Liu, H.Y. - Fu, J.Y. Strajniski vestnik - Journal of Mechanical Engineering 66(2020)12, 697-708 (20000 N), so both of the two controllers have achieved good track tracking performance. However, RBFSMESO has better trajectory performance than the PID does due to the former's good model and external interference compensation capability. The maximum trajectory tracking error is no more than 0.006 mm, while the maximum tracking error of the PID controller reaches 0.42 mm. At the same time, by comparing the tracking error curves of the two, it can be seen that the PID controller's tracking error significantly increases when constant external interference is added at t> 5 s, while the RBFSMESO's tracking error increases instantaneously, but with the accurate observation and compensation of external interference of the disturbance observer, its tracking error converges quickly. When t> 15 s, the PID controller has a large tracking error deformation after sinusoidal external interference is used, and external interference has no basically influence on the RBFSMESO controller, which further verifies the good anti-interference capability of the RBFSMESO controller. Fig. 7. shows the observation performance and observation error of extended state observer (ESO) for external interference in RBFSMESO. When external disturbance of constant value is added at t> 5 s, the ESO observation error increases instantaneously. However, when external disturbance value tends to be stable, the observation error rapidly converges. The instantaneous maximum disturbance observation error is no more than 50 N. When sinusoidal external interference is applied, the observation error increases due to the bandwidth limitation of the observer, but the maximum tracking error is still no more than 120 N, which further verifies the good observation performance of the observer. Meanwhile, by comparing Figs. 6 and 7, it can be seen that the trajectory tracking error of RBFSMESO corresponds 30 Time [s] Fig. 7. The observation performance and observation error of ESO Fig. 8. PID trajectory-tracking performance and tracking error RBF Neural Network Sliding Mode Control Method Based on Backstepping for an Electro-hydraulic Actuator 703 Strajniski vestnik - Journal of Mechanical Engineering 66(2020)12, 697-708 to the disturbance observation error of ESO, which verifies the importance of accurate disturbance observation and compensation to the controller performance. 2. High-speed tracking and strong external interference: Expected curve of trajectory tracking: x1d = 0.015 sin(2nt)[1 - exp(-10t3)], (43) and external load interference: fax = 0 t < 5 s fd2 = 120000{1 - exp[-0.1x (t - 5)3]} t > 5 s fd3 = 30000sin(2^t) t > 15 s (44) In the case of high-speed tracking and strong external interference, the trajectory-tracking performance and tracking error of the PID and RBFSMESO controllers are shown in Figs. 8 and 9. By comparing with Figs. 5 and 6, it can be seen that with the increase of interference, the trajectory-tracking performance of the two controllers deteriorates to a certain extent, and the maximum tracking error reaches 2.8 mm and 0.0087 mm respectively. It can be found that RBFSMESO's trajectory tracking-performance is far better than PID due to its good anti-interference ability. Fig. 10 shows the observation performance and observation error of ESO to external interference. The comparison with Fig. 7 shows that with the increase of external interference, the observation error also increases to a certain extent. However, compared with the maximum external interference of 150000 N, the maximum observation error is 666 N, which still shows the good observation performance of the observer. Fig. 9. Step response curve of backstepping method Fig. 10. The observation performance and observation error of ESO 704 Li, W.D. - Shi, G.L. - Zhao, C. - Hongyu Liu, H.Y. - Fu, J.Y. Strajniski vestnik - Journal of Mechanical Engineering 66(2020)12, 697-708 4 EXPERIMENTAL VERIFICATION An experiment was carried out on the semi-physical experimental platform to verify further the feasibility of the strategy. The semi-physical experimental platform consists of an inertial load, a set of force loading equipment, a set of position measurement systems and a set of actuators testers. The inertial load is used to simulate the magnitude of the inertial force and the frictional damping. The force loading device can apply the predetermined force and can be used to simulate the disturbing force. The position measurement system can measure the position response of the load in real time. Because the connection is not completely stiff, the measured value will be different from the displacement output of the actuator itself in dynamic motion. The actuator tester is used for the instruction issuing of electro-hydraulic actuators and the detection and data analysis of each parameter. The displacement, velocity, and differential load pressure were collected via sensors [31]. 1. Low-speed tracking condition Given the following trajectory tracking expectation curve: x1d = 0.015sin(0.159x2nt)[1 -exp(-10t3)]. (45) The PID and RBFSMESO controllers are used to verify the trajectory tracking of the curve given above. The parameters of the controller are specified in Section 3.2. The trajectory-tracking performance and tracking error curves of the two controllers are shown in Figs. 11 and Fig. 12, respectively. The maximum trajectory tracking errors of the two controllers are 1.181 mm and 0.4567 mm, respectively. Combined with the above simulation results, it can be seen that RBFSMESO further improves the trajectory tracking accuracy of the system through the function of RBF neural network and sliding mode control on Fig. 11. PID trajectory-tracking performance and tracking error Fig. 12. RBFSMESO trajectory-tracking performance and tracking error RBF Neural Network Sliding Mode Control Method Based on Backstepping for an Electro-hydraulic Actuator 705 Strajniski vestnik - Journal of Mechanical Engineering 66(2020)12, 697-708 the basis of introducing disturbance observation and compensation. However, the experimental error of the two controllers is greater than the simulation error, because the simulation model ignores some modelling error terms of the system, including the servo valve dynamics, the friction modelling error of the actuator and other unmodeled high-frequency dynamics, which leads to the increase of the trajectory tracking error of the controller. 2. High speed tracking condition Expected curve of trajectory tracking: x1d = 0.015 sm(2^)[1 - exp(-10t3)]. (46) PID and RBFSMESO controllers are still used to verify the trajectory tracking of the curve given above. The trajectory-tracking performance and tracking error of the PID and RBFSMESO controllers under high-speed tracking are shown in Figs 13 and 14, respectively. With the increase of the expected trajectory motion frequency, the tracking performance of the two controllers declined, and the maximum tracking error of the two controllers reached 3.045 mm and 1.033 mm, respectively. The main reason is that with the increase of dynamic frequency, the dynamic response of the servo valve has a great impact on the control performance. When the controller is designed, the servo valve dynamic is simplified to a simple proportional link, which is suitable for low-speed tracking. However, under high-speed tracking, its dynamic response cannot be ignored, so it should be modelled as a first-order or second-order model for design, which is also worthy of further improvement. 5 CONCLUSIONS In this article, ESO is introduced, then an RBF neural network sliding mode control method based Fig. 13. PID trajectory-tracking performance and tracking error Fig. 14. RBFSMESO trajectory-tracking performance and tracking error 706 Li, W.D. - Shi, G.L. - Zhao, C. - Hongyu Liu, H.Y. - Fu, J.Y. Strajniski vestnik - Journal of Mechanical Engineering 66(2020)12, 697-708 on backstepping for an electro-hydraulic actuator is proposed. 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D0l:10.5545/sv-jme.2020.6714 Original Scientific Paper Received for review: 2020-04-10 Received revised form: 2020-06-22 Accepted for publication: 2020-07-06 Influence of Wear and Tool Geometry on the Chatter, Cutting Force, and Surface Integrity of TB6 Titanium Alloy with Solid Carbide Cutters of Different Geometry Jianyong Liu1 - Jianfei Sun2 3 4 - Uzair Khaleeq uz Zaman5* - Wuyi Chen23 1 AECC Shenyang Liming Aero-Engine Co., Ltd., China 2 Beihang University, School of Mechanical Engineering and Automation, China 3 Collaborative Innovation Center of Advanced Aero-Engine, China 4 Beijing Engineering Technological Research Center of High-efficient & Green CNC Machining Process and Equipment, China 5 National University of Sciences and Technology, College of Electrical and Mechanical Engineering, Pakistan In this paper, vibration-free milling cutters (variable helix (VH) and variable pitch (VP) end mills) and standard (SD) end mills are used to machine TB6 (Ti-10V-2Fe-3Al) titanium alloy in order to study the influence of wear and geometric structure parameters of milling cutters on chatter, cutting force and surface integrity of machined surfaces. The results of the tests show that the wear of milling cutters has a significant influence on the chatter, cutting force, roughness, residual stress, and microhardness. Geometric structure parameters of milling cutters also have a clear impact on both chatter and cutting force. Also, chatter and cutting force have significant effects on roughness and residual stress, which are in turn affected by tool geometric structure parameters, separately. Keywords: chatter, cutting force, surface integrity, TB6 Titanium alloy, wear Highlights • An analysis was performed to study the effects of wear, chatter, and cutting force on the surface integrity of TB6 titanium alloy when machined with solid carbide cutters of different geometry. • Surface integrity has an important influence on surface quality and is influenced by many factors, such as vibration, cutting parameters, tool wear and tool geometric structure parameters. • As per experimental results, the wear of milling cutters has a significant influence on the chatter, cutting force, roughness, residual stress, and microhardness. • Chatter and cutting force have significant effects on roughness and residual stress. 0 INTRODUCTION Titanium alloy has been widely used in various fields, such as aerospace, automotive, medical equipment, etc., because of its underlying properties, including high specific strength, excellent corrosion resistance, and good fatigue resistance [1] and [2]. TB6 titanium alloy is a near p phase titanium alloy, which is denser and presents high strength at low operating temperatures [3]. TB6 titanium alloy is primarily used in aircraft structural components (aircraft fuselage, wing, landing gear, and helicopter rotor parts) [4]. However, some characteristics of TB6 titanium alloy, such as low thermal conductivity, low elastic modulus, and high chemical activity, can cause high cutting temperatures, large cutting forces, shorter tool life, low metal removal rates, and poor work surface integrity [5] and [6]. The improvement of surface integrity of TB6 titanium is, therefore, a challengeable subject in the area of manufacturing due to its low machinability. Surface integrity is commonly defined as "the topographical, mechanical, chemical and metallurgical state of a machined surface and its relationship to functional performance [7]. It consists of factors such as surface roughness, work hardening, residual stress, microstructure, etc. Moreover, surface integrity is an important performance characteristic of the machined surface quality, thereby determining the functionality and fatigue life of the critical structure components [8]. Surface integrity is impacted by many factors, including vibration, cutting force, tool wear, and similar. Pimenov et al. [9] studied the effect of the relative position on machined surface roughness with a face-milling cutter. The results proved that the relative position of the milling cutter and workpiece affected the cutting force, vibration and, surface roughness. In 1907, chatter, as a machining phenomenon, was first introduced by [10], after which it has been extensively studied [11]. The chatter during machining has a significant effect on surface integrity, tool wear, machine damage, etc. It can not only reduce the surface quality of the workpiece and increase tool wear but can also cause damage to the machine [12]. Several methods have been used in the literature *Corr. Author's Address: College of Electrical and Mechanical Engineering, National University of Sciences and Technology, Islamabad, 44000, Pakistan, uzair.khaleeq@ceme.nust.edu.pk 709 Strojniški vestnik - Journal of Mechanical Engineering 66(2020)12, 709-723 to reduce vibration in the cutting process, and the most widely used technique is to optimize cutting conditions based on the stability lobe diagram [13] and [14]. Since the cutting parameters are restricted by the stability lobe diagram, only cutting parameters within a particular range can effectively suppress chatter. Reducing the spindle speed will reduce vibration but at the cost of lower material removal rate and a decrease in machining efficiency. If the cutting speed is increased, the cutting temperature will also increase, thereby affecting the surface quality and tool wear because of high chemical activity and low thermal conductivity of TB6 titanium alloy. Furthermore, the variable helix (VH) and variable pitch (VP) milling cutters can effectively suppress chatter at low cutting speeds [15] and [16] making their use an economical method to reduce the vibration and improve the cutting stability without changing the cutting speed by using the vibration-free milling cutters. Many researchers have indicated that the geometric structure parameters of milling cutters had significant influence on the quality of machined surfaces. Particularly when the variable helix (VH) and variable pitch (VP) tools were involved, the investigations were usually carried out to enhance stability and suppress regenerative chatter in order to improve the milling quality. Ott et al. [17] studied the mechanical vibrations in milling with non-uniform pitch and variable helix tools. The results showed that the distributed delay in variable helix tools led to a stabilization of the cutting process. Moreover, Sims et al. [18] analysed the influence of radial immersion on the stability of variable pitch or helix milling tools. The cyclic fold bifurcations were found to exist for both at lower radial immersions. Also, Wang et al. [19] investigated the chatter prediction for variable pitch and variable helix milling and found them to be effective for suppressing chatter. Stability lobes were studied by Altintas et al. [20] for variable pitch milling cutters and designed an optimal tooth spacing to increase the chatter-free depth of cuts. Li et al. [21] also investigated the effect of a face-milling cutter with a variable pitch on vibration. Furthermore, Niu et al. [22] improved the generalized Runge-Kutta method by considering runout to analyse the milling process stability with variable pitch and variable helix milling cutters. However, the published literature on the VH and VP milling cutters seldom analysed the chatter with tool wear and rarely emphasized surface quality indexes other than chatter induced surface texture. Furthermore, cutting forces have a great influence on workpiece precision, surface quality, cutting system vibration, cutting power, and tool life, especially when the milling cutters have a complex geometry in machining [23]. Cutting forces are affected by several factors such as tool geometry, workpiece material properties, cutting conditions, etc. [24]. Huang et al. [25] investigated the cutting forces by analysing the time domain and the frequency domain in TC4 titanium alloy milling. The results showed that the radial forces of VP milling cutter were minimum compared with standard (SD) and VH milling cutters. Moreover, tool wear has a significant impact on the surface integrity of titanium alloy. The influence of tool wear on roughness is related to the tool wear state. Liang and Liu [26] investigated the effect of tool flank wear on surface integrity; the results showed that the roughness first increased and then decreased with the tool flank wear increasing from 0 mm to 0.3 mm. This behaviour was caused by different wear states of tool flank face. However, Yang et al. [27] reported contradictory results, in which the machined surface roughness of TB6 presented an increasing trend with the increase of tool flank wear. Tool wear also had significant influence on residual stress and microhardness. Residual stresses beneath the machined surface became much more compressive and penetrated to a deeper depth with the worn tools [28]. Nevertheless, some other researchers obtained the contrary results, with the tool flank wear increasing, and the surface residual stresses becoming more tensile due to the increased temperature [29]. The hardness of the machined surfaces produced with worn tools was harder, and the affected layer depth was larger than those produced with new tools under the same cutting parameters [30] and [31]. Consequently, based on the literature reviewed, the reasons affecting the surface integrity of titanium alloy are complex, and a gap in research exists about the influence of wear, chatter, and cutting force on surface integrity with vibration-free milling cutters. The study in this paper investigates the surface integrity of the titanium alloy TB6 machined by VH, VP, and SD milling cutters, and aims to provide reference data for the effects of wear, tool geometric structure parameter, chatter, and cutting force on surface integrity. 1 EXPERIMENTAL DESIGN Tool wear and surface integrity tests were designed for the subject purpose. First, a wear test was carried out to observe the flank wear of the milling cutters. Then the cutting test of surface integrity was carried out after the flank wear of the three types of milling 710 Liu, J. - Sun, J. - Zaman, U.K. - Chen, W. Strojniški vestnik - Journal of Mechanical Engineering 66(2020)12, 709-723 cutters reached the blunt standard. Cutting force and vibration were recorded during the surface integrity test. VH, VP, and SD milling cutters were used in the cutting tests of wear and surface integrity. K44UF cemented carbide was used as tool material for the three types of milling cutters, which were not coated. The geometric structure parameters of the three types of milling cutters are shown in Table 1. Table 1. Geometric structure parameters of three types of milling cutters Milling No. of Radius of Helix Tooth cutter flutes tool [mm] angle [deg] pitch [deg] VP 4 5 45 86 and 94 VH 4 5 41 and 45 85 and 95 SD 4 5 45 90 The three types of milling cutters had the same geometric structure parameters except for the helix angle and tooth pitch. The cross-section of three types of milling cutters is shown in Fig. 1. Fig. 1. Cross-section of three types of milling cutters; the lengths are in mm The two cutting tests were then carried out under dry milling conditions, and the workpieces were machined by side milling and up milling. The size of workpiece used in the two cutting tests was (70 x 57.5 x 35) mm. The workpiece was machined with an NC machine tool of type BV75. The material of the workpiece used in all tests was TB6 titanium alloy, which is one of the p-phase titanium alloys with high strength, high toughness, and machined after the forging process. The nominal chemical composition of TB6 titanium alloy is shown in Table 2, and the mechanical properties at room temperature of the TB6 alloy are shown in Table 3 [32]. Wear of the milling cutters was observed using the Dino-Lite AM7013MZT microscope and expressed by the width of flank wear of milling cutters. The blunt standards of the three types of milling cutters were 0.1 mm, 0.2 mm, and 0.3 mm, respectively. The width of flank (the main cutting edge) wear of milling cutters was measured when the wear of cutting edge was close to the blunt standard. Then the mean value of four cutting edges was taken as the wear extent of the milling cutter. Table 2. Chemical composition of TB6 (mass fraction, %) Element wt. [%] Element wt. [%] Al 3.15 O 0.10 V 10.5 N 0.03 Fe 2.1 H 0.002 C 0.02 Ti Balance Table 3. Mechanical properties of TB6 at room temperature Mechanical properties Value Tensile strength [MPa] 1105 Yield strength [MPa] 1035 Elastic modulus [GPa] 106 Elongation [ %] 4—10 Hardness [HV] > 375 Fracture toughness [MPa Vm] > 60 Cutting force was measured with a piezoelectric force sensor of type CL-YD-3310, and the acceleration was measured with a piezoelectric acceleration transducer of type VM45-1QZH. The radial rigidity of milling cutter was low, and since the radial cutting force was more suitable for analysing the vibration of the cutting system than the tangential and axial cutting forces [33] and [34], it was studied in this paper. The vibration and cutting force measurement devices of the low rigidity cutting test are shown in Fig. 2. Furthermore, the modes of the cutting system were measured with hammering tests. The first-order modes of these two test systems are shown in Table 4. Fig. 2. Measuring equipment of vibration and cutting force of low rigidity cutting test Influence of Wear and Tool Geometry on the Chatter, Cutting Force, and Surface Integrity of TB6 Titanium Alloy with Solid Carbide Cutters of Different Geometry 711 Strojniški vestnik - Journal of Mechanical Engineering 66(2020)12, 709-723 Table 4. First-order modes of cutting systems Cutting system Workpiece-dynamometer-worktable [Hz] Tool-spindle [Hz] Mode 469 216 During the cutting process, the depth of cut and the width of cut were kept constant, and the overhang of the three types of milling cutters was 40 mm each. The cutting parameters used in the two cutting tests are shown in Table 5. Table 5. Cutting parameters of wear and surface integrity cutting tests Test no. Cutting speed [m/min] Feed per tooth [mm/z] Depth of cut [mm] Width of cut [mm] Milling cutter 1 30 0.06 5 0.5 2 30 0.06 5 0.5 SD 3 30 0.06 5 0.5 4 30 0.06 5 0.5 5 30 0.06 5 0.5 VH 6 30 0.06 5 0.5 7 30 0.06 5 0.5 8 30 0.06 5 0.5 VP 9 30 0.06 5 0.5 Surface roughness was measured by using the roughness measuring instrument of type TIME3220. The roughness of the workpiece surface was measured along the feed direction. The sampling length of roughness measurement was 0.8 mm, and the evaluation length was 4 mm. Samples were cleaned using ethanol solution via ultrasonic cleaning before taking measurements. Nine samples were measured in this experiment, and the average roughness was obtained by measuring roughness five times for each sample. The residual stress was measured with the Prism system, produced by Stresstech Oy. The hole-drilling technique was used for residual stress measurement in which, by removing a volume of material, residual stresses were released locally, and the stress equilibrium in the part was changed. The stresses in the remaining material rebalanced and the surface distorted slightly, especially near the hole. The Prism system measured surface distortion optically using a laser that was diffusely reflected from the sample surface. For the nine samples measured, the distances below the surface where the holes were drilled were 0.01 mm, 0.02 mm, 0.03 mm, 0.04 mm, 0.05 mm, 0.060 mm, 0.07 mm and 0.080 mm. Nano-hardness at different depths on the cross-section of the machined surface was measured using a Nano Indentationer G200 system. Nine tests were performed to obtain the nano hardness. The three indentations performed below the machined surface in each test were 5 ^m, 50 ^m, and 100 ^m, respectively (as shown in Fig. 3). The penetration depths were realized using diamond Berkovich tip after precise focus on the cross-section. The penetration depths for each indentation below the surface were 1^m. Machined surface Test 1-9 Indentation 1 Indentation 2 ' indentations 5tim î Î f" 50 jim 1 inntim Fig. 3. Distribution of indentations in nanometer indentation measurement Microstructures were observed by using the three-dimensional (3D) laser microscope LEXT 0LS4100. First, the nine samples were embedded to three specimens, and each specimen included three samples. Then, the samples were abraded and polished. Finally, the samples were etched for about 20 seconds using a corrodent matched by HF : HNO3 : H2O = 1 : 4 : 45. 2 RESULTS AND DISCUSSION 2.1 Wear The wear extents were determined by measuring the flank wear of the VH, VP, and SD milling cutters. First, the flank wear of each cutting edge of the three types of milling cutters was measured for five times and the average of the wear of the four cutting edges was taken to obtain the wear extent of each milling cutter. The blunt standard of the three types of milling cutters was the same for all, and the blunt standard was 0.1 mm, 0.2 mm, and 0.3 mm. When cutting speed was set to 30 m/min, and the feed was fixed at 0.06 mm, the wear extent of each milling cutter changed with time; the changes observed are shown in Fig. 4. It can be deduced from Fig. 4 that the wear of VH, VP, and SD milling cutters was at the normal wear stage. When the wear extents were close to 0.1 mm, 0.2 mm, and 0.3 mm, the wear of VP milling cutter was the fastest, followed by VH milling cutter, and SD milling cutter. Therefore, SD milling cutter 712 Liu, J. - Sun, J. - Zaman, U.K. - Chen, W. Strojniski vestnik - Journal of Mechanical Engineering 66(2020)12, 709-723 Fig. 4. Wear extent varied with time of three types of milling cutters has the best wear resistance, and the VP milling cutter has the worst wear resistance. As shown in Fig. 5, when the wear of milling cutter was close to the blunt standard, which was 0.1 mm or 0.2 mm, the wear measuring errors of teeth of VP and VH milling cutters were big and that of SD milling cutter were small. This phenomenon is due to the changes of helix and pitch causing noneven wear of teeth of vibration-free milling cutters, thereby increasing the wear measuring errors. With the increase of wear, the wear-measuring errors of VP and VH milling cutter edges had a decreasing tendency, especially when the wear was close to the blunt standard of 0.3 mm. Also, the wear measuring 0.16 0 14 0.12 £ J. 0.10 c ro | 0.08 0 06 0.04 a) 0.241 0.230 220.210 200.190,18- b) 0320 0.3150.310 _0 305-| 0.300 0 295-J 2 0 2900.2850 2800 275- 025- 0 20- | 0 15-c a a o 10 0 05 1st 2nd 3rd Teeth of SD milling cutter 0.00- 4th 1st 2nd 3rd Blades of SD milling cutter 4th 030 0 25 | 0 20 S |0 15 0 10 0.05 1st 2nd 3rd 4 th Blades of SD milling cutter 0 36-, 0 34- „0.32- E E "0.30- tn o> 2 0.280 26- 0.24- 1st 2nd 3rd 4th Teeth of VH milling cutter 0.18 0.16 0 14 Fo 12 E 10 |0 08 0.06 0.04 - 002 0.30 0.28 0.26 0.24 E* 0 22 E — 0.20 I 0 1S * 0.16 0.14 0.12 0.10 1st 2nd 3rd 4th Blades of VH milling cutter 0.36 0.34 •§■0 32 50 30 0.28 0.26 1st 2nd 3rd 4th Blades of VH milling cutter 1st 2nd 3rd 4th Teeth of VP milling cutter 1st 2nd 3rd 4th Blades of VP milling cutter 1st 2nd 3rd 4lh Blades of VP milling cutter Fig. 5. Wear measuring errors of teeth of three types of milling cutters; a) VB = 0.1 mm, b) VB = 0.2 mm, and c) VB = 0.3 mm Influence of Wear and Tool Geometry on the Chatter, Cutting Force, and Surface Integrity of TB6 Titanium Alloy with Solid Carbide Cutters of Different Geometry 713 c) Strojniški vestnik - Journal of Mechanical Engineering 66(2020)12, 709-723 errors of VP and VH milling cutter edges decreased significantly, thereby causing reductions in the nonuniform wear caused by the cutter geometry (helix angle and pitch). The wear-measuring errors of the three types of milling cutters are shown in Table 6, which shows that the error between the measured wear of the three types of milling cutters and the blunt standard was between 0.1 % and 5.1 %, and the measured wear was close to the blunt standard. Table 6. Measuring errors of wear Cutting speed Blunt standard Wear Error Milling [m/min] [mm] [mm] [%] cutter 0.1 0.1048 4.8 30 0.2 0.2082 4.1 SD 0.3 0.2949 5.1 0.1 0.1014 1.4 30 0.2 0.1965 1.75 VH 0.3 0.3121 4.03 0.1 0.0986 1.4 30 0.2 0.2014 0.7 VP 0.3 0.3004 0.1 The changes of pitch and helix angles of vibration-free milling cutters induced the cutting force to change, leading to faster wear of VP and VH milling cutters compared to wear rate of SD milling cutter. This fact further resulted in even wear of SD milling cutter and non-even wear of VP and VH milling cutter [35] thereby increasing the wear measuring the difficulty of VP and VH milling cutters. 2.2 Chatter 2.2.1 Resonance Analysis The vibration frequency is mainly related to the spindle rotation frequency (SRF) and tooth pass frequency (TPF). SRF and TPF are used to analyse whether chatter occurs during the cutting process. SRF is defined as shown in Eq. (1). kn 1000Av SRF = — =-, 60 60nD (1) where, n is spindle speed [r/min], v is cutting line speed [m/min], D is milling cutter diameter [mm], and k is the coefficient with values 1, 2, 3,... Moreover, TPF is defined as shown in Eq. (2). kn TPF = N —, 60 (2) where, N is the number of teeth of the milling cutter. According to Eqs. (1) and (2), when the cutting speed was 30 m/min, SRF = 15.9 Hz, and TPF = 63.6 Hz. Moreover, when the occurred frequency was different from multiples of SRF and TPF, the chatter appeared. Figs. 6 to 8 showed the vibration spectrum of SD, VH and VP milling cutters when cutting speed was 30 m/min, the feed was 0.06 mm, and the wear of the three types of milling cutters reached the blunt standard which was 0.1 mm, 0.2 mm and 0.3 mm, respectively. Ac max in these figures represented the maximum chatter amplitude. The first mode of tool-spindle was 216 Hz, and the first mode of workpiece-dynamometer-worktable was 469 Hz (see Table 4). Since the maximum amplitude frequency was far from the two modes in the low rigidity cutting test (see Figs. 5 to 7), there was no resonance generation at the frequency of maximum amplitude, but there were amplitudes that had approximate frequency near the two modes, and the amplitudes were small. Therefore, small resonance occurred in the cutting test of surface integrity. 2.2.2 Influence of Wear on Chatter As shown in Fig. 9, the maximum chatter amplitude of SD, VH and VP milling cutters increased with the increase of wear. When the wear extent was 0.1 mm, the chatter of SD milling cutter was the largest, followed by VH milling cutter, and VP milling cutter, and there was a small difference of chatter between the VH and VP milling cutters. When the wear extent was 0.2 mm and 0.3 mm, the chatter of SD milling cutter was the largest, followed by VP milling cutter, and VH milling cutter. 2.3 Cutting Force For studying the average cutting force, it can be seen from Fig. 10 that the cutting force of the three types of milling cutters increased when the wear increased from 0.1 mm to 0.3 mm. When the wear extent was 0.1 mm and 0.2 mm, the cutting force of VP milling cutter was the largest, followed by SD milling cutter, and VH milling cutter. When the wear extent was 0.3 mm, the cutting force of SD milling cutter was the largest, followed by VP milling cutter, and VH milling cutter. The increase of the contact area between the flank face of milling cutter and the workpiece led to the increase of the extrusion pressure. However, due to the different geometric structure parameters of SD, VH, and VP milling cutters, the extrusion pressure 714 Liu, J. - Sun, J. - Zaman, U.K. - Chen, W. Strojniški vestnik - Journal of Mechanical Engineering 66(2020)12, 709-723 Frequency[ a) b) c) Fig. 6. Acceleration spectrum of standard milling cutter; a) VB = 0.1, V = 30 m/min, f = 0.06 mm/z; b) VB = 0.2, V = 30 m/min, f = 0.06 mm/z; c) VB = 0.3, V = 30 m/min, f = 0.06 mm/z Frequency [kHz] i-requency ikmzj Frwtjwty p a) b) c) Fig. 7. Acceleration spectrum of variable helix milling cutter a) VB = 0.1, V = 30m/min, f = 0.06 mm/z; b) VB = 0.2, V = 30 m/min, f = 0.06 mm/z; c) VB = 0.3, V = 30 m/min, f = 0.06 mm/z a) b) c) Fig. 8. Acceleration of variable pitch milling cutter a) VB = 0.1, V = 30m/min, f = 0.06 mm/z; b) VB = 0.2, V = 30 m/min, f = 0.06 mm/z; c) VB = 0.3, V = 30 m/min, f = 0.06 mm/z was different at the same wear. Compared with SD milling cutter, the actual feed per tooth of VP milling cutter was not consistent, because it was affected by the change of tooth pitch and the superposition of given feed, which also became the reason behind the change in cutting force. For the VH milling cutter, compared with SD milling cutter, except that the feed of each tooth pitch was inconsistent, which caused the different actual feed, the change of helix angle led to the change of working rake angle, the actual feed and working rake lead subject to change of cutting force. The wear of milling cutter led to the decrease of radial cutting depth which reduced the cutting force. These reasons caused the cutting forces of the three types of milling cutters to turn out to be different. 715 Influence of Wear and Tool Geometry on the Chatter, Cutting Force, and Surface Integrity of TB6 Titanium Alloy with Solid Carbide Cutters of Different Geometry 33 Strojniški vestnik - Journal of Mechanical Engineering 66(2020)12, 709-723 0.2 Wear extent (mm) Fig. 9. Chatter varied with wear extent of three types of milling cutters 0.2 Wear extent [mm] Fig. 10. Cutting force varied with wear extent of three types of milling cutters 2.4 Surface Integrity 2.4.1 Surface Roughness 2.4.1.1 Influence of Wear on Roughness The roughness of machined surface produced by VH, VP and SD milling cutters was measured five times, and the mean roughness and standard deviation are shown in Fig. 11. The standard deviations indicated that the discrete degree of measured roughness was small, and the measurement of roughness had high precision. It can also be seen from Fig. 11 that the roughness of machined surface produced by three types of milling cutters increased with the increase in wear. a) b) 0.2 cj Wear extent [mm] Fig. 11. Roughness of machined surface with three types of milling cutters at different wear, a) VH, b) VP, and c) SD 2.4.1.2 Influence of Chatter on Roughness Roughness was affected by the chatter. When the roughness produced by SD, VH, and VP milling 716 Liu, J. - Sun, J. - Zaman, U.K. - Chen, W. Strojniski vestnik - Journal of Mechanical Engineering 66(2020)12, 709-723 cutters were compared (see Fig. 12), the roughness caused by SD milling cutter was the largest, followed by VP milling cutter, and VH milling cutter. The reason for this is that the chatter had a significant influence on roughness. Chatter can increase the radial displacement of the cutter teeth in cutting, Fig. 12. Influence of geometric structure of milling cutters on surface roughness increasing radial cutting depth, thereby increasing the roughness. Compared with SD milling cutter, when the wear extent was 0.2 mm and 0.3 mm, the effect of suppressing chatter of VH milling cutter was the best, followed by VP milling cutter. Although the chatter of the VP milling cutter was the smallest when the wear extent was 0.1 mm, there was little difference of chatter between the VP and the VH milling cutters (as shown in Fig. 9). The roughness was, therefore, affected not only by the chatter but also by the tool geometric structure parameters. The change of tooth pitch affected the actual feed causing the change of helix angle, which in turn caused the working rake angle, thereby affecting the roughness. For these reasons, the roughness produced by VH milling cutters was the smallest. 2.4.2 Residual Stress The generation process of residual stress is complex. High temperature, pressure, strain, and strain rate are produced in the contact area of milling cutter and workpiece, and uneven elastoplastic deformation occurs in the cutting area. Generally, residual stress is caused by uneven elastoplastic deformation caused a) b) c) Fig. 13. Residual stress of machined surface with VH milling cutter; a) VB = 0.1 mm, b) VB = 0.2 mm, and c) VB = 0.3 mm a) b) c) Fig. 14. Residual stress of machined surface with VP milling cutter; a) VB = 0.1 mm, b) VB = 0.2 mm, and c) VB = 0.3 mm Influence of Wear and Tool Geometry on the Chatter, Cutting Force, and Surface Integrity of TB6 Titanium Alloy with Solid Carbide Cutters of Different Geometry 717 Strojniški vestnik - Journal of Mechanical Engineering 66(2020)12, 709-723 by cutting force and cutting heat. Therefore, residual stress is the result of the interaction of cutting parameters, tool wear, tool geometric structure parameters and several other factors in the cutting process. In the present work, the resultant residual stresses exhibited irregular distribution (as shown in Figs. 13 to 15). It was found that the residual stress had large fluctuation, which might be caused by the large grain size and uneven plastic deformation of TB6 titanium alloy [36]. It can also be seen that the residual compressive stress tended to be the main stress with the increase of wear. Moreover, Fig. 16 showed the maximum residual stress of machined surface with VH, VP, and SD milling cutters resulted when the wear extent was 0.1 mm to 0.3 mm. It can be further observed that the maximum residual tensile stress and the maximum residual compressive stress increased with the increase of wear. When the milling cutter was worn, the contact area between the flank face of milling cutter and the workpiece surface increased, thereby increasing the extrusion pressure on the workpiece, which increased residual compressive stress. At the same time, the wear of milling cutter increased the friction between flank face and the workpiece, causing an increase in the friction heat, thereby causing the increase of residual tensile stress. The workpiece surface is also affected by both the cutting force and cutting heat in the cutting process, and residual stress is the result of the combined action of the two factors. The depth range of 0 mm to 0.04 mm from the machined surface was close to the machined surface of the workpiece. The heat dissipated rapidly in this depth range, and the extrusion effect of cutting force was obvious. Therefore, the residual stress in this depth range was dominated by compressive stress, and the maximum residual compressive stress appeared in this depth range. Furthermore, the depth range of 0.04 mm to 0.08 mm was far from the machined surface. Due to the poor thermal conductivity of TB6 titanium alloy, the thermal stress had an obvious effect on the material in this depth range. Therefore, the residual stress within this depth range was dominated by tensile stress, and the maximum residual tensile stress appeared within this depth range. 2.4.3 Microhardness Nano-hardness was obtained via three steps: loading of sample, dwelling, and unloading. At the start, a) b) c) Fig. 15. Residual stress of machined surface with SD milling cutter; a) VB = 0.1 mm, b) VB = 0.2 mm, and c) VB = 0.3 mm 0,2 Wear extent [mm] a) 0,2 Wear extent {mm) b) 0.2 Wear extent [mm] c) Fig. 16. Maximum residual stress of machined surface with three types of milling cutters; a) VH, b) VP, and c) SD 718 Liu, J. - Sun, J. - Zaman, U.K. - Chen, W. Strojniški vestnik - Journal of Mechanical Engineering 66(2020)12, 709-723 loading of the sample was performed from the depth of 0 nm to 1000 nm; dwelling was executed from 1000 nm to 1024 nm, and, finally, unloading was done from 1024 nm to 769 nm. The obtained loading/unloading curve is shown in Fig. 17. 150 300 450 600 750 900 Displacement into surface [nm] Fig. 17. Loading/unloading curve As shown in Fig. 18, the sub-surface of nine machined samples were divided into three regions according to the positions of intentions. These were characterized as the hard region, soft region, and near the bulk material region. Fig. 19 showed the nano hardness in the hard region, soft region, and near bulk material region with the wear from 0.1 mm to 0.3 mm. The hardness of the material in the hard region was the greatest, followed by the near bulk material region, and the hardness of the material in the soft region was the smallest. It can also be observed that the geometric structure parameters of the three types of milling cutters had little influence on the hardness distribution beneath the machined surface. Hard region Idention y Soft region Near bulk 4 < material region 50 um Idention 3—' f t '/jjflfl 200 um Bulk material hardness, value=5.73GPa lQOtym Fig. 18. Regions below the machined surface During the processing, the titanium elements in the titanium alloy interacted with oxygen elements and nitrogen elements in the air to form titanium oxide and titanium nitride, which made the machined surface brittle [37] and increased the microhardness of the hardening zone. The extrusion pressure on the workpiece surface was also an important reason for the increase of hardness in the hardening zone. Furthermore, the thermal conductivity of TB6 titanium alloy was low, which caused the cutting heat to be too late to transfer in cutting titanium alloy, 6 1 6.0 acq s « 5.8 ai I 5.7. n O nj z 5.5 5.4 -vp-j—so - Bulfc material hardne». valued .''-'">•'■! Hard 'i'-','; * - _Softfeflion Near bulk maternal regttfl 5 50 100 Depth below the machined surface [mm] 6.2 I as.o £ 5.8 o 5.6 c ra Z 5,4 5.2 - VM—VP—*— SD Bu& material hordness, vatoe=575GPa Hard region Near bulk matenal region i. 1 \ 1 \ i \ 6.6 6.4 £6.2 O «8.0 If) £5.8 "D 15.6 0 1 5.4 5.2 - 5.0 5 50 100 Depth below the machined surface [mm] -VP—t-SO BüBl material r>ai 28; AW represents the cross-sectional area of wick structure calculated by Eq. (14). Thus, the F can be calculated by Eq. (15). 2f- r 2 K = £ u fRe s = - n(W + Wb )8 + 5 I -n A n (W + Wb ) 2n(dv +5) (10) (11) 1 v y 2 Plough-extrusion Forming for Making Micro-groove Heat Pipes on Hydrostatic Thrust Bearings of Heavy Machinery 727 Strojniški vestnik - Journal of Mechanical Engineering 66(2020)12, 724-735 ru = 2 A, / C,, (W + Wb )s - W ) + 452 (12) (13) AW + S) "^(d^l (14) F = (( +V(W - W )2 + 452 )2 n (W + Wb ) S3p,h (15) fg Fig. 5. Drag coefficient of laminar flow in trapezoidal micro-groove [28] 4) Fv is the friction coefficient of the steam flow according to the flow conditions of the steam, which is related to the steam Reynolds number and the Mach number. The steam Reynolds number Rev can be calculated with Eq. (16). Re, = 2rhvQ Av /Jvh (16) ■fg where, rhv= djl for micro heat pipe with a circular steam chamber, then Eq. (16) can be expressed as Eq. (17). Re = - 4Q ndv ßvh (17) fg The Mach number Mv can be calculated by Eq. (18). 728 Q Mv =- ,- A Pvh/gyJ Yvrvt — 4Q ndvpv h fg FTM' . (18) Then the parameter Fv can be calculated by Eqs. (19) to (22) according to different situations. a. If Rev < 2300 and Mv< 0.2 F, = 128^ AvrÜv Pvhfg ndl Pvhfg (19) b. If Rev < 2300 and Mv> 0.2 F. =- AvrL Pvhfg 128^ 1 + ^ M2 nd1pvh fg 1 + M2 -12 >2 (20) c. If Rev > 2300 and Mv < 0.2 F = 0.038»v AhvPvh 0.608»v ndvpvh fg 2rhvQ v Avhfg »v, 4Q ndvhfg /uv (21) d. If Rev > 2300 and Mv > 0.2 F = 0.038^, Ar Pvhfg 0.608^ ndvpvh fg 2rhvQ V Ahfg Vv, 4Q ndvhfg /uv i + >Vzl m 2 - i + Ml .(22) Thus, the capillary heat transfer limit of the trapezoidal micro-groove heat pipe can be calculated by Eq. (23). 4aS Qca W V(W - Wb )2 + 4S2 = -plgdv cos^iplglsin^ ( +J(W - Wb )2 + 4S2 ) ) n(W + Wb ) Slp,hA - + F -. (23) leff Thus, when the temperature of working fluid and the effective length of the micro heat pipe are constant, the capillary limit is influenced by the diameter of the steam chamber, the top and bottom width of the trapezoidal micro-groove, the depth and number of the grooves and the flow conditions of the steam. 46 Li, X. - Yu, Z. - Li, X. - Li, W. - Zou, T. r,, = hl 2 2 2 4 Strojniski vestnik - Journal of Mechanical Engineering 66(2020)12, 724-735 2 PLOUGH-EXTRUSION FORMING USING MULTI-TOOTH TOOL Plough-extrusion forming of metal uses the special plough-extrusion multi-tooth tool on the machine tool to plastically deform a part of the metal on a metal wall and build fins with many micro-grooves. Thus, plough-extrusion forming is a metal-processing technology with the characteristics of both plough and extrusion. Although it uses the lathe machining, and its tool is similar to the turning tool, the forming mechanism is different from turning. Plough-extrusion forming has both characteristics of plough and extrusion deformation. Therefore, it can be called plough-extrusion forming. Turning makes the required shape by cutting the extra metal from the material. However, plough-extrusion forming produces plastic deformation and slicing via the effect of extruding to form grooves and fins without cutting. Thus, the fin it produces is much deeper than what the turning tool does. tool rvi Fig. 6. The diagram of plough-extrusion forming for making fin structure The workpiece material in this research is copper T2, and the tool material is bearing steel GCr15 or high-speed steel W18Cr4V The workpiece is mounted on the machine tool table, as shown in Fig. 6. As mentioned before, the process of plough-extrusion forming is different from turning on common machine tools. The turning process forms the required shape by cutting the extra metal, and the removed metal materials often become chips. However, in plough-extrusion forming, a special multi-tooth tool is used to plough, crush, and tear to form sub-structures and micro-structures of fins, which are superimposed on the macro-structure. There are no chips or only a few cutting materials during machining. For making sub-structures and micro-structures of fins, the entire plough-extrusion forming process can be divided into four stages: cut-in, extrusion, fin forming, finishing. The cut-in process uses the multi-tooth tool to cut the copper surface with the front blade and forms the opening of groove, then the surface metal is ploughed by pressure in front of the multi-tooth tool, and forcing the metal to start diverting to both sides. In the extrusion step, after the surface layer of copper is cut in, the two extrusion surfaces of the multi-tooth tool begin to extrude the metal. With the continuous deepening of the multi-tooth tool, when the applied stress of the multi-tooth tool reaches the yield limit of the copper material, the copper begins to undergo plastic deformation. The copper is then forced to flow in the direction of minimum resistance after hindered by metal resistance. In other words, the copper flows outward along the normal direction of the tool pressing surface, and the metal bulging occurs continuously on both sides of the pressing surface to form the first fin. The fin-forming process is relatively complicated, and under the combined forces of friction and extrusion, the first formed fins begin to crack to form a jagged structure. In the case in which the multi-tooth tool continues to plough and extrude, the top of each fin is re-cracked to form a periodic serrated fin structure. In the finishing step, after ploughing and extruding to form a relatively complete fin structure, the secondary extrusion surface of the multi-tooth tool will still extrude and repair on the fins formed in the previous step. At this time, a small amount of metal will undergo minor plastic deformation, and flow along the secondary extrusion surface of the multi-tooth tool to make the fin structure grow upward and form the sub-structure and micro-structure superimposed on the macrostructure. Through repeated cycles of these above four processes, the sub-structures and micro-structures of fin are continuously produced and stacked. 3 DESIGN AND MANUFACTURE OF MULTI-TOOTH TOOL 3.1 Material Selection for Multi-tooth Tool High-speed steel is especially suitable for the manufacture of complicated tool with high thermal stability, high strength, a certain degree of hardness and wear resistance. It is easy to be shaped into a sharp cutting edge. The ploughing tool is more complex, so bearing steel material GCr15 or high-speed steel material W18Cr4V are used. The bar diameter is chosen to be 20 mm. Plough-extrusion Forming for Making Micro-groove Heat Pipes on Hydrostatic Thrust Bearings of Heavy Machinery 729 Strojniški vestnik - Journal of Mechanical Engineering 66(2020)12, 724-735 3.2 Design of Multi-tooth Tool The parameters of six designed multi-tooth tools are shown in Table 1 (the parameters in parentheses are derived from the geometric calculation, not from direct design). The cross-section of a single tooth of the multi-tooth tools is an isosceles trapezoid, and the angle between the two sides of a single tooth is 120°. The tool length is 10 mm, the tooth number is 50, 55 or 60, and the depth of groove is 0.2 mm or 0.25 mm. As the number and depth of tooth increase, the plough-extrusion force is also increased, and the micro-chips will become more numerous, which accelerates the wear of the tool. 3.3 Manufacture of Multi-tooth Tool The manufacturing process of the multi-tooth tool is as follows: 1) The raw material is cut into bars of 10 mm length. 2) A centre hole of ®5 mm is drilled on a ®20 mm bar for tool mounting. 3) Heat treatment and quenching steps are shown in Figs. 7a and b. There are two common kinds of materials for making multi-tooth tool for different needs, W18Cr4V6 and GCr15. For high-speed steel W18Cr4V6, it is first quenched (oil cooling) between 1200 °C and 1300 °C, and then tempered 3 times at 550 °C for 1 hour each time; For bearing steel GCr15, it is first quenched between 800 °C and 860 °C (oil cold), and then tempered 1 time between 150 °C and 170 °C for 2 to 3 hours. 4) The process of wire electrical discharge machining for a multi-tooth tool is shown in Figs. 7c, d and e. In order to improve the efficiency of wire electrical discharge machining, the five treated bars are put on the positioner together to perform cutting, and the use of positioner can help to locate the central axis of the inner hole during the wire electrical discharge machining. A low-speed is taken to minimize the error because the wire may break if the speed is fast. If the wire breaks, the micro-grooves of the Fig. 7. Heat treatment and wire electrical discharge machining in multi-tooth tool producing; a) heating, b) quenching, c) clamping, d) workpiece localization, e) wire electrical discharge machining, f) formed multi-tooth tool 730 Li, X. - Yu, Z. - Li, X. - Li, W. - Zou, T. Strojniski vestnik - Journal of Mechanical Engineering 66(2020)12, 724-735 tool cannot be machined correctly due to the positioning problem. In other words, it is not able to correspond with the point and form the correct number of micro-grooves and that will also waste the raw material of the tool. Therefore, the voltage and current of wire electrical discharge machining should be controlled within a reasonable range. In addition, the cutting wire should be as close as possible to the tool, but do not approach the material; otherwise, it will cause greater error. 5) Grinding. One end of the multi-tooth tool is sharpened, and reasonable parameters are obtained, such as the anterior horn, the dorsal horn, the rake face and the rear face. 6) Cleaning. After cleaning, the final formed multi-tooth tool is acquired as shown in Fig. 7f. 4 EXPERIMENTS OF PLOUGH-EXTRUSION FORMING AND THE RESULTS The plough-extrusion forming experiment was performed on the XK5032 CNC vertical lifting table milling machine. The whole procedure was as follows: 1) A plate was clamped on the worktable as shown in Fig. 8a. 2) A 3-mm-deep hole was drilled using the centre drill to effectively reduce the error from shaking when the drill meets copper block as shown in Fig. 8b. 3) The hole was drilled and reamed to an aperture of ®10 mm as shown in Fig. 8c; 4) The multi-tooth tool was mounted on the pull rod to perform plough-extrusion forming as shown in Fig. 8d. The position of the horizontal plane of the multi-tooth tool was fine-tuned, and the chamfer of the multi-tooth tool was used to guide the positioning. Then, let the pull rod directly go down without rotating for plough-extrusion forming, and the illustration of the process was shown in Fig. 8e. Finally, the formed inner micro-grooves on the plate were shown in Fig. 8f. Fig. 9 shows the comparison between wire electrical discharge machining and plough-extrusion forming for making the structure of the heat pipe. Figs. 9b and e were acquired by the digital camera equipment and it was found that the wire electrical discharge machining method did not make the weight and volume of the micro-groove larger when compared with the plough-extrusion forming method, and the surface of groove was smooth without the fin structure shown in Fig. 9f. However, the surface of micro-groove made by plough-extrusion forming was rough as shown in Figs. 9d and e. The surface roughness of the groove can increase the area of the surface, thereby increasing the wettability of the working fluid and improving the heat transfer performance of the flat heat pipe. Fig. 9e shows the enlarged view of micro-grooves manufactured by plough-extrusion forming, where the formed grooves have better height, Fig. 8. Plough-extrusion forming for micro-grooves by multi-tooth tool; illustration of; a) clamping, b) central positioning, c) of drilling, d) clamping multi-tooth tool, e) plough-extrusion forming, f) formed inner micro-grooves Plough-extrusion Forming for Making Micro-groove Heat Pipes on Hydrostatic Thrust Bearings of Heavy Machinery 731 Strojniški vestnik - Journal of Mechanical Engineering 66(2020)12, 724-735 depth, and width than the ones made by wire electrical discharge machining. The plough-extrusion forming also made a secondary groove structure superimposed on the main groove structure, as shown in Fig. 9c to improve the roughness. Furthermore, the wire electrical discharge machining method is more complicated and costs more than the plough-extrusion forming method. The plough-extrusion forming can even be carried out on a simple self-made cutting platform. Thus, the plough- extrusion forming method has an advantage of making micro-grooves providing higher capillary force at a lower manufacturing cost. For the experiment, six multi-tooth tools with different designed geometric sizes (Table 1) were adopted to compare their forming performance in plough-extrusion procedure. All these tools were made of bearing steel GCr15.The intercepted formed micro-groove of the heat pipe was magnified and focused with a 12-million-pixel digital camera device, (a) . - m * » * (b) m * * * g * m # f ^ 10 tr f ;HE3 Fig. 9. The comparison between wire electrical discharge machining and plough-extrusion forming for making the structure of heat pipe; a) micro-grooves manufactured by wire electrical discharge machining, b) enlarged view of micro-grooves manufactured by wire electrical discharge machining, c) the structures of main and secondary grooves from enlarged view of micro-grooves manufactured by plough-extrusion forming, d) micro-grooves manufactured by plough-extrusion forming, e) enlarged view of micro-grooves manufactured by plough-extrusion forming, f) the structures of fin and groove from enlarged view of micro-grooves manufactured by plough-extrusion forming 732 Li, X. - Yu, Z. - Li, X. - Li, W. - Zou, T. Strojniski vestnik - Journal of Mechanical Engineering 66(2020)12, 724-735 Table 1. Design parameters of multi-tooth tool Number of tool 1 2 3 4 5 6 Tooth number 50 50 55 55 60 60 Angle for each tooth [°] 7.20 7.20 6.54 6.54 6.00 6.00 Tooth top width [mm] 0.25 0.25 0.22 0.22 0.20 0.20 Tooth bottom width [mm] (0.1973) (0.1842) (0.1673) (0.1542) (0.1473) (0.1342) Tooth fin width [mm] (0.2527) (0.2527) (0.2365) (0.2365) (0.2189) (0.2189) Tooth depth [mm] 0.2 0.25 0.2 0.25 0.2 0.25 and then the photographed surface was compared and analysed. According to the comparison, it was found that the multi-tooth tool with 10 mm length, 55 tooth number, 0.25 mm tooth depth and 0.22 mm tooth top width, which corresponds to No. 4 in Table 1, could achieve the best performance in these six designed tools. Different forming speed was also tested to determine the best setting. The testing plough-extrusion forming speed was increased from 30 mm/min to 70 mm/min using the tool with best performance, and the corresponding length, height, width and spacing of the formed fin structure were recorded in Fig. 10. It can be found that with the increase of plough-extrusion forming speed, the fin height, the fin width and the spacing between two fins changed little, and they tended to be stable. Furthermore, the fin length varied slightly with speed. Thus, 50 mm/min is taken to be the most suitable forming speed for acquiring the largest fin length in this experiment. In order to verify the effectiveness of the theoretical model in guiding the actual manufacturing, the verification test was carried out in experiments. The heat pipes used for the verification were made of copper, and were formed by the multi-tooth tool with the parameters of No. 5 item in Table 1 according to the method described in this research. The number of the micro-grooves is 60, the depth of the micro-grooves is 0.20 mm, and the bottom width of the micro-grooves is 0.15 mm, the width of the top side of the micro-grooves is 0.20 mm, and the micro-grooves are evenly distributed in the circumferential direction. The formed micro-groove copper tube shell was Fig. 10. The relationship between the plough-extrusion forming speed and the structure parameters of serrated fins Plough-extrusion Forming for Making Micro-groove Heat Pipes on Hydrostatic Thrust Bearings of Heavy Machinery 733 Strojniški vestnik - Journal of Mechanical Engineering 66(2020)12, 724-735 pumped to a negative pressure of 1.3^(10-1 to 10-4) Pa and then filled with an appropriate amount of working fluid, which was water in the test, so that the capillary porous material of the wick was filled with liquid and then sealed to form a micro-groove heat pipe. During the test, the heat pipe is placed horizontally, the temperature of the evaporation section is kept at 60 °C and the condensation section is cooled with water to below 30 °C. Fig. 11. Example of experimental system for heat transfer test The effective length of the micro heat pipe is 235 mm, of which the length of the evaporation section is 50 mm, the length of the adiabatic section is 85 mm, and the length of the condensation section is 100 mm. The relevant parameters of water at 60 °C are shown as below: M = 18.0156, pv= 1.99919x104 Pa, hgg = 2.3584x106 kJ/kg, pj = 983.28 kg/m3, pv= 0.1302 kg/m3, = 4.63x10-4 N-s/m2, ^v= 1.06x10« N-s/m2, k = 0.653 W/(m-K), a = 6.607x10-2 N/m. According to the theoretical model described in Section 1, the theoretical parameters of the heat pipe can be calculated as below: its viscosity limit is 3.01x105 W, sound velocity limit is 1.31x103 W, carrying limit is 223 W, capillary limit is 79.3 W, and boiling limit is 6.82x103 W. Because its maximum heat transfer is determined by the capillary limit, the maximum theoretical heat transfer of a micro-groove heat pipe with these parameters is 79.3 W. Ten micro-groove heat pipes formed in the experiments with the same parameters as the theoretical sample were tested by the specific heat transfer testing system as shown in Fig. 11. Their heat transfers were 80.3 W, 78.6 W, 79.6 W, 79.2 W, 78.9 W, 79.3 W, 79.7 W, 78.9 W, 80.1 W and 78.8 W. The average value was 79.4 W, which had a deviation of 0.1 W compared with the maximum theoretical heat transfer. Using the scanning electron microscope (SEM) to analyse the cross-section of the micro-groove heat pipe, it was found that due to the influence of various factors in the processing, as well as the influence of plastic deformation flow and partial recovery, there was a certain error in the actual micro-groove. However, the error between the experimental value and the theoretical value was within an acceptable range, and the model in Section 1 could be considered correct. 5 CONCLUSION The heat transfer performance of the heat pipe mainly depends on the structure of the inner wall wick. The plough-extrusion forming method proposed in this paper can quickly make micro-grooves in the inner hole of flat. However, there are micro-chips in the machining process. In order to ensure smooth discharge of the chips, vertical top-down processing is adopted. According to image analysis of forming experimental results, the multi-tooth tool with 10 mm length, 55 tooth number, 0.25 mm tooth depth and 0.22 mm tooth top width was chosen to be the most suitable forming tool and the 50 mm/min forming speed was taken to achieve optimization. During the manufacturing procedure, if the forming speed is too fast, the pull rod may be damaged. However, if the forming speed is too slow, the effect of ploughing is not obvious, and it is a waste of time. 6 ACKNOWLEDGEMENTS This project is supported by the National Natural Science Foundation of China (grant no. 81701087), Natural Science Foundation of Fujian Province in China (2017J01681) and the Outstanding Youth Fund of Fujian Agriculture and Forestry University (XJQ201820). 7 REFERENCES [1] Yu, M., Yu, X., Zheng, X., Jiang, H. (2019). Thermal-fluid-solid coupling deformation of hydrostatic thrust bearing friction pairs. Industrial Lubrication and Tribology, vol. 71, no. 3, p. 467-473, D0l:10.1108/ILT-07-2018-0262. [2] Wang, D.Q., Yan, B.H., Chen, J.Y. (2020). The opportunities and challenges of micro heat piped cooled reactor system with high efficiency energy conversion units. Annals of Nuclear Energy, vol. 149, art. ID 107808, D0I:10.1016/j. anucene.2020.107808. [3] Liang, L., Diao, Y.H., Zhao, Y.H., Wang, Z.Y., Bai, F.W. (2020). 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Sonic Limitations and Startup Problems of Heat Pipes. NASA Tech Brief. [25] Faghri, A., Thomas, S. (1989). Performance characteristics of a concentric annular heat pipe: Part I—experimental prediction and analysis of the capillary limit. Journal of Heat Transfer, vol. 111 no. 4, p. 844-850, DOI:10.1115/1.3250795. [26] Chi, S.W. (1976). Heat Pipe Theory and Practice. Hemisphere Publishing Corp., Washington, DC. [27] Tien, C.L., Sun, K.H. (1971). Minimum meniscus radius of heat pipe wicking materials. International Journal of Heat and Mass Transfer, vol. 14, no. 11, p. 1853-1855, DOI:10.1016/0017-9310(71)90052-4. [28] Kays, W.M. (2012). Convective Heat and Mass Transfer Tata McGraw-Hill Education, New York. Plough-extrusion Forming for Making Micro-groove Heat Pipes on Hydrostatic Thrust Bearings of Heavy Machinery 735 Strojniški vestnik - Journal of Mechanical Engineering 66(2020)12, 736-751 © 2020 Journal of Mechanical Engineering. All rights reserved. D0I:10.5545/sv-jme.2020.6929 Original Scientific Paper Received for review: 2020-09-10 Received revised form: 2020-11-26 Accepted for publication: 2020-12-02 Experimental Identification of Optimized Process Parameters for FSW of AZ91C Mg Alloy Using Quadratic Regression Models C. Satheesh.1 - P. Sevvel2 - R. Senthil Kumar2 1 Madha Institute of Engineering and Technology, Department of Mechanical Engineering, India 2 S.A. Engineering College, Department of Mechanical Engineering, India This experimental work aims to devise and establish quadratic regression equations, including various input criteria of a friction stir welding (FSW) technique to predict and determine the responses during the fabrication of AZ91C Mg alloy joints. The input process parameters taken into consideration include the traversing speed of the tool, the speed of rotation of the tool, its pin profile (geometry) and the axial force. A five-level, 4 four-factor composite design (of central nature) was applied, and response surface methodology (RSM) was used to formulate quadratic regression models, to develop 3D response surface charts, and to anticipate the responses for various mechanical properties. The generated quadratic mathematical model was tested and validated using the technique of analysis of variance. Validation experimental trial results outlined in the form of scatter diagrams revealed precedented coincidence with that of the generated models. The AZ91C Mg alloy joints obtained using the tool having taper cylindrical pin geometry employed at 1045 rpm, 1.5 mm/s traversing speed, under the exertion of an axial load of 4.87 kN was found to exhibit improved mechanical properties. Keywords: AZ91C Mg alloy, quadratic regression model, friction stir welding, tool pin geometry, tool rotational speed, ultimate tensile strength, tool traversing speed, axial force Highlights • An experimental endeavour was made to predict the ideal and optimized combination of input parameters of FSW technique to achieve maximized mechanical properties including yield, tensile strengths and percentage of elongation, through the employment of a statistically and mathematically advanced technique namely response surface methodology (RSM). • The established quadratic regression model was efficient in determining the responses of the FSWed AZ91C Mg allot joints in the range of +10 % of their actual experimental values at a 95 % level of confidence. • 3D response surface charts illustrating the influence of input parameters on the mechanical properties and the direct impact plots representing the individual impact of these parameters on mechanical strength were generated and analysed in detail. • AZ91C Mg alloy joints obtained using a tool having taper cylindrical pin geometry, employed at 1045 rpm, along a 1.5 mm/s tool travelling speed, under the exertion of a 4.87 kN axial force, were found to be free from flaws and naturally exhibited exceptional mechanical properties, especially higher values of ultimate tensile strength • Six AZ91C Mg alloy joints were fabricated, employing the generated optimized input parameters. All the six joints were examined, and the observations revealed that they are completely free from defects, which proves their exact conformity with the generated values. 0 INTRODUCTION Among the lightweight category materials, alloys of magnesium, especially AZ91C Mg alloy, are employed for an ample range of aerospace, automobile, and shipbuilding sectors, given their excellent specific strength, recyclability, higher values of specific stiffness, outstanding machinability, lower density values etc. [1] and [2]. However, at the same time, being a hexagonally closed packed structured metal, the slip structures of alloys of Mg are constrained concerning plastic deformation. Therefore, improving the ductility of Mg alloys (especially AZ91C) is an essential criterion for increasing the usage and applicability of these alloys to various industrial sectors [3] and [4]. It is a proven fact that the ductility and strength of the metals possessing hexagonal closed packed structures can be enhanced through the refinement of their grain structures. It is foreseen through various research works that the state of super-plasticity will occur in these alloys if their grain size is reduced drastically [5]. AZ91C is a very common and popular Mg alloy, which is employed in nearly 85 % to 91 % of all the Mg cast components, due to its unique and excellent cast-ability properties, low cost, high levels of resistance to corrosion, etc. At the same time, its lower values of elongation property (arising due to the existence of coarse grains in its microstructure) have to be improved to make it applicable to various industrial sectors [6] to [8].With the objective of widening the application areas of this AZ91C Mg alloys to several industrial sectors, various research initiatives [9] to [13] have been carried out broadly, primarily focusing on the improvement of the ductility of the AZ91C Magnesium alloys through the refinement of their grain structures. For instance, the Mg alloy (Mg-9Al-1Zn) was experimentally proved to exhibit higher 736 *Corr. Author's Address: S.A. Engineering College, Department of Mechanical Engineering Tharapakkam, Chennai - 600 122, India, sevvel_ready@yahoo.co.in Strojniški vestnik - Journal of Mechanical Engineering 66(2020)12, 736-751 levels of superplasticity at reduced temperatures, due to the employment of differential speed rolling process by Kim et al. [11]. This experimental investigation also revealed that, apart from the size of the grains, the microformability was also greatly influenced by the transition of superplastic flow to non-superplastic flow. Compared to various plastic deformation improvement techniques, friction stir welding was very effective in fabricating high-quality aluminium similar and dissimilar joints, as the joining process takes place without their respective melting points. This fact was well proven by Khodabakhshi et al. [14] through his experimental investigation carried out during the friction stir processing of an Al-Mg alloy (AA5052) sheet, employing the combination of Monte Carlo based simulation and modelling based on computational fluid dynamics modelling / Monte Carlo simulation to speculate the refinement of grains during that joining process. The parameters of friction stir welding (FSW) (i.e., speed of tool rotation and velocity of traverse) were taken as input parameters for formulating a numerical model for understanding the dynamic recrystallization of grains. The results reveals that the peak temperature was the preeminent factor in controlling the fabricated weldment's grain microstructure. Recent research [15] to [23] carried out with respect to friction stir welding has demonstrated that improvement has taken place in the mechanical properties of Al alloy joints, due to the refinement of the grain structures and modifications in the microstructure of the welded regions. For example, the experimental work by Yuan et al. [17] investigated in detail the impacts of heat treatment (post weld) on the stability and formability of the friction stir welded tubes. The investigational results proved that the welded tubes' formability properties had been improved, because of the refinement of grains. It was proved through this investigation that the size of the subgrains and recipitates were controlled by this heat treatment process (carried out after the welding) and the stability of the microstructure is controlled by this heat treatment, which indirectly dominates the formability and strain hardenability of the friction stir welded tubes. The practical possibility of employing friction stir welding for fabricating joints under the water was experimentally recorded by Derazkola et al. [19] during the joining of Al alloy (AA5083) sheets with Steel (AISI A441) sheets by submerging them under the water at three different temperatures: 50 °C, 25 °C, and 0 °C. It was experimentally proven that, by reducing the cooling medium's temperature, the intermixing of the dissimilar metals was crumbled, leading to the generation of the intermetallic compound in the interface region. It was also found that the nugget zone's hardness was decreased due to the increased cooling ability of the submerged water medium. The precipitate distribution, size of the grains, and microhardness of the several regions of friction stir-welded T87 Al alloy was investigated by Liang et al. [20]. It was observed during this investigation that recrystallization occurred dynamically in the centre of nugget zone, and the grains in these regions are smaller in size when compared to that of the grains in the other weld zones. Moreover, the size of the precipitates was very much smaller in the heat zone when compared to that of the parent metal. Another critical task in producing sound quality joint for any welding technique is directly related to the employment of ideal and optimal input parameters [24] and [25]. The specimens welded under different combinations of process parameters derived using the conventional techniques require an inspection to decide whether they meet the expected standards or not. These conventionally derived parameters are likely to fabricate joints nearer to the customer expectations; what is missed is the desirable combination of the optimized process parameters. In other words, alternative, better combinations of welding process parameters always exist, which can be employed, only if they can be determined. In this experimental investigation, an endeavour has been made to predict the ideal and optimized combination of input parameters of FSW technique to achieve maximized mechanical properties including yield, tensile strengths and percentage of elongation, through the employment of a statistically and mathematically advanced technique: response surface methodology. 1 TECHNIQUE OF EXPERIMENTATION 1.1 Process Parameters Identification The material of our investigation is AZ91C Mg alloy and its chemical composition includes: Al - 9.15 %; Cu - 0.08 %; Zn - 0.859 %; Si - 0.119%; Mn - 0.3 %; Fe and Ni - both less than 0.01 % and Mg as the remaining percentage. The alloy's tensile strength was 287 MPa, along with a 191 MPa Yield strength and percentage of elongation being 8.74. In an FSW technique, the parameters'(i.e., speed of rotation of the FSW tool) force applied from upwards, angle of tilt of the FSW tool, tool traverse velocity along weld line, shape of the tool pin profile, tool shoulder dimension Experimental Identification of Optimized Process Parameters for FSW of AZ91C Mg Alloy Using Quadratic Regression Models 737 Strojniski vestnik - Journal of Mechanical Engineering 66(2020)12, 736-751 Table 1. Description of employed process parameters, their levels and factors adopted in this experimental work Sample Parameters of FSW Levels No. - 2 - 1 0 1 2 1 Geometry of tool pin P Straight square Straight cylindrical Threaded cylindrical Taper cylindrical Taper threaded cylindrical 2 Rotational speed [rpm] NR 500 650 800 950 1100 3 Tool traversing speed [mm/s] nt 0.5 0.75 1.00 1.25 1.5 4 Axial force [kN] F 2 3 4 5 6 and etc. are found to play determine the quality and the result of the entire process. As it is practically impossible to identify and understand the influence of all the process parameters of FSW technique, the parameters that were reported and recorded (by several researchers) as most dominant and which seem to have a significant part in inducing the strength of the fabricated joints are presented in Table 1, along with their different levels. Because of the broad spectrum of influence by the abovementioned process parameters during the fabrication of joints using FSW technique, a central composite design matrix comprising 31 trail runs, along with five levels, four factors of significance designed using the technique of response surface methodology. From Table 1, it can be understood that the maximum and minimum limit for these significant factors were assumed to be +2 and -2, respectively. 1.2 Description of Employed FSW Tools In this experimental investigation, the flat plates of cast AZ91C Mg alloy (5 mm thickness) were joined by FSW technique by employing tools with five different pin geometries (as mentioned in Table 1); straight square, taper cylindrical, straight cylindrical, taper threaded cylindrical and threaded cylindrical. The photographic illustration of these above-mentioned tools with five different pin geometries are shown in Fig. 1. FSW tools with five different pin geometries employed in this experimental investigation Fig. 1. All these above-mentioned friction stir welding tools were fabricated using the M35 grade high-speed steel. 1.3 Experimental Design Matrix and Investigational Runs Table 2 describes in detail the 31 sets of the experimental, investigational positions and their respective responses. The 31 experimental runs of the coded positions and conditions constitute a bisected replication of 16 factorial design (24) together with eight start (axial) points and points of the centre being 7. The points of the centre are constituted by the entire process parameters of the FSW placed at the levels of middle (i.e., 0). Simultaneously, the axial points are constituted by the combinations of individual FSW process parameters, either at their highest (+2) value or at their least (-2) value together with other four process parameters at their respective intermediate levels. Hence, these foresaid 31 investigational runs permit the evaluation of the quadratic equations and double-way interactive and collective impacts of these input parameters of FSW on the various responses (i.e., yield, tensile strength and percentage of elongation). As indicated by the matrix of design of the response surface methodology, a total of 31 investigational runs were carried out employing the FSW technique and AZ91C Mg alloy FSW joints (Experimental run no. 31) were fabricated. During this process, the direction of joining was carried out in such a way that the travel direction of tool was assimilated to the direction of rolling of the AZ91C Mg alloy flat plates and each joint was obtained using a single pass. The photographic view of these fabricated AZ91C Mg alloy joints are shown in Fig. 2. In this experimental and mathematical analysis, the material being employed for fabricating the tools for joining AZ80A Mg alloy flat plates is M35 grade high-speed steel (HSS). Tools with three distinctive pin geometries as shown in Fig. 1 are employed in this experimental work for fabricating the AZ80A Mg alloy joints using FSW technique. 738 Satheesh, C. - Sevvel, P. - Senthil Kumar, R. Strojniški vestnik - Journal of Mechanical Engineering 66(2020)12, 736-751 Table 2. Description of Design Matrix and their corresponding responses Experimental runs Parameters Experimental responses Experimental runs Parameters Experimental responses P Nr nt F UTS YS Percentage of elongation P Nr nt F UTS YS Percentage of elongation 1 -1 1 1 1 190 93 3.71 17 0 2 0 0 226 130 4.25 2 1 1 1 -1 230 136 4.73 18 0 0 0 0 219 121 3.51 3 0 0 -2 0 210 117 3.70 19 -1 -1 1 -1 181 82 3.05 4 1 1 -1 -1 228 134 4.82 20 2 0 0 0 207 115 4.27 5 -1 1 -1 1 185 90 3.65 21 -2 0 0 0 148 51 2.12 6 0 0 0 0 218 122 3.51 22 1 -1 -1 1 220 128 4.58 7 -1 -1 -1 1 182 84 3.02 23 -1 1 -1 -1 187 90 3.64 8 -1 -1 -1 -1 178 81 2.94 24 0 0 0 0 217 122 3.5 9 -1 -1 1 1 184 87 3.05 25 1 -1 1 1 223 130 4.83 10 1 -1 1 -1 215 127 4.49 26 0 0 0 0 219 121 3.51 11 0 0 2 0 213 119 3.91 27 0 0 0 0 217 120 3.52 12 0 0 0 2 209 112 3.94 28 1 -1 -1 -1 218 126 4.69 13 1 1 1 1 238 143 5.05 29 0 0 0 0 217 122 3.5 14 0 -2 0 0 175 89 3.25 30 0 0 0 0 215 120 3.49 15 1 1 -1 1 229 135 4.98 31 -1 1 1 -1 186 89 3.5 16 0 0 0 -2 198 113 3.81 2 MATHEMATICAL MODEL 2.1 Correlation Equations Response surface methodology (RSM) is an assortment of rational, statistical and arithmetical approaches, applied for evaluating and investigating the problems in which a single response is influenced by a considerable amount of individual variables [26] and [27]. The basic purpose and intention of employing RSM to an investigational problem are to determine and decide the exact arrangement and sequence of the parameters of a process, which will provide the best and excellent response value. Apart from this, the response surface methodology presents a model of regression, which portrays the connection between the responses and parameters of a process. This connection can be employed to anticipate the Fig. 2. AZ91C Mg alloy joints fabricated using the design matrix based experimental investigational runs) photos taken during the 1st part of experimental runs and b) photos taken during the 2nd part of experimental runs Experimental Identification of Optimized Process Parameters for FSW of AZ91C Mg Alloy Using Quadratic Regression Models 739 Strojniski vestnik - Journal of Mechanical Engineering 66(2020)12, 736-751 response, whenever the parameters of an investigated process are diversified within the selected boundaries [28]. These regression models analytically describe the experimental surfaces, when the plot has been drawn for responses versus any two parameters of an investigational process. These generated plots make it feasible to understand the connections between the parameters and responses for a given experiment [29]. In our investigation, the response functions namely, yield strength, tensile strength (ultimate) and percentage of elongation of the fabricated joints are functions of input parameters namely, FSW tool pin geometry (P), rotating speed of tool (NR), traversing speed of tool (NT) and force acting from the upward direction (F). They are expressed as: UTS = 217.43 + (18.58 P) + (7.25NR) + (1.08NT) + (2.08-F) - (8.92P2) - (3.17-NR2) - (0.42-N^) - (2.42-F2) + (1.62PNR) + (0.12PNT) + (0.62PF) - (0.37-AR-F) + (1.13NrF), (1) Yield = 121.14 + (20.46 P) + (6.13NR) + (0.96-Nt) + (0.96F) - (8.88P2) - (2.25NR2) - (0.13NT2) - (1.50-F2) + (0.56^ PNR) + (0.44^ PNT) + (0.06- PF) + (0.31NRNT) - (0.06NRF> + (0.81^NT"F), (2) Elongation = 3.506 + (0.663- P) + (0.226^ NR) + (0.021NT) + (0.053 F) - (0.013P2) + (0.126NR2) + (0.139NT2) + (0.157-F2) - (0.09 b PNR) - (0.002- P NT) - (0.002-P-Nt) + (0.026- P F) - (0.018NrNt) + (0.024NRF> + (0.046NrF). (3) 2.2 Confirming the Competence of Established Model The competency of the generated and proposed mathematical model is analysed by employing the ANOVA approach, and the outcomes of the 2nd order response surface model were found to coincide with the output models of ANOVA (analysis of variance). Results of ANOVA for the response functions (i.e., tensile strength (ultimate) and yield strength of the fabricated joints) are indicated in the Tables 3 and 4, respectively. Table 3. Results of ANOVA for response functions of tensile strength (ultimate) Source DF Seq SS Contribution [%] Adj SS Adj MS F-Value P-Value Model 14 12244.8 93.91 12244.8 874.63 17.61 0.000 Linear 4 9682.0 74.25 9682.0 2420.50 48.73 0.000 P 1 8288.2 63.56 8288.2 8288.17 166.87 0.000 Nr 1 1261.5 9.67 1261.5 1261.50 25.40 0.000 nt 1 28.2 0.22 28.2 28.17 0.57 0.462 F 1 104.2 0.80 104.2 104.17 2.10 0.167 Square 2485.3 19.06 2485.3 621.32 12.51 0.000 PP 1 2072.1 15.89 2275.1 2275.08 45.80 0.000 NrNr 1 244.9 1.88 287.3 287.29 5.78 0.029 nt-nt 1 0.8 0.01 5.0 5.04 0.10 0.754 FF 1 167.4 1.28 167.4 167.42 3.37 0.085 2-Way Interaction 77.5 0.59 77.5 12.92 0.26 0.948 PNr 1 42.2 0.32 42.2 42.25 0.85 0.370 pnt 1 0.2 0.00 0.2 0.25 0.01 0.944 PF 1 6.2 0.05 6.2 6.25 0.13 0.727 nrnt 1 6.2 0.05 6.2 6.25 0.13 0.727 NrF 1 2.2 0.02 2.2 2.25 0.05 0.834 Nr F 1 20.3 0.16 20.3 20.25 0.41 0.532 Error 16 794.7 6.09 794.7 49.67 Lack-of-Fit 10 783.0 6.00 783.0 78.30 40.10 0.000 Pure Error 6 11.7 0.09 11.7 1.95 Total 30 13039.5 100.00 Model Summary: S - 7.04767; R - Sq: 93.91 % 740 Satheesh, C. - Sevvel, P. - Senthil Kumar, R. Strojniški vestnik - Journal of Mechanical Engineering 66(2020)12, 736-751 Table 4. Results of ANOVA for response functions of yield strength Source DF Seq SS Contribution [%] AdjSS Adj MS F-Value P-Value Model 14 13347.8 95.96 13347.8 953.4 27.14 0.000 Linear 4 10989.5 79.00 10989.5 2747.4 78.20 0.000 P 1 10045.0 72.21 10045.0 10045.0 285.93 0.000 NR 1 900.4 6.47 900.4 900.4 25.63 0.000 nt 1 22.0 0.16 22.0 22.0 0.63 0.440 F 1 22.0 0.16 22.0 22.0 0.63 0.440 Square 2338.0 16.81 2338.0 584.5 16.64 0.000 P 1 2144.8 15.42 2254.6 2254.6 64.18 0.000 Nr 1 128.4 0.92 145.3 145.3 4.14 0.059 nt 1 0.0 0.00 0.5 0.5 0.01 0.908 F 1 64.7 0.47 64.7 64.7 1.84 0.194 2-Way Interaction 20.4 0.15 20.4 3.4 0.10 0.996 PNr 1 5.1 0.04 5.1 5.1 0.14 0.709 pnt 1 3.1 0.02 3.1 3.1 0.09 0.772 PF 1 0.1 0.00 0.1 0.1 0.00 0.967 nrnt 1 1.6 0.01 1.6 1.6 0.04 0.836 NrF 1 0.1 0.00 0.1 0.1 0.00 0.967 Nr F 1 10.6 0.08 10.6 10.6 0.30 0.591 Error 16 562.1 4.04 562.1 35.1 Lack-of-Fit 10 557.2 4.01 557.2 55.7 68.84 0.000 Pure Error 6 4.9 0.03 4.9 0.8 Total 30 13909.9 100.00 Model Summary: S-5.92720; fl-Sq: 95.96% If the obtained value of Fratio of the developed model is lower than the definitive Fratio value at a named 95 % to 96 % certainty level, the established and predicted model will be competent with the level of confidence. The value of probability > F for the three established models is less than the value of 0.05 (i.e., 95 % level of confidence), which indirectly proves that the developed model is satisfactory. The predicted values versus the actual experimental values for the responses of the developed regression equations are graphically in Fig. 3, which reveals the presence of an exceptional interrelationship between the predicted value of the generated responses and the actual experimental values. In these graphs, it can be seen that the distributed and scattered plots are very much nearer to the 45° line, which shows the flawless fitness of the predicted mathematical relationship [30]. 3 OUTCOMES, VALIDATION, AND DELIBERATION 3.1 Micro and Macrostructural Analysis Fig. 4 illustrates the macrographical image of a cross-section of the sound quality welded specimen taken horizontally to the FSW tool's travelling direction. The nugget zone of this welded specimen is found to be completely free from various defects, such as porosity, tunnel voids, excessive flash, kissing bond defects, defective tightness, onion skin microstructures, crack-similar root flaws, etc. [31] to [33]. Fig. 5 describes in detail the microstructure of the parent metal and different zones (heat-affected portion (zone) (HAZ), thermos mechanically affected portion (zone) (TMAT) and nugget zone (NZ) of the defect-free welded specimen. It can be visualized from these microstructures that AZ91C Mg alloy contains large coarse sized grains, unevenly distributed over the entire region. However, at the same time, from the micrographs of the various regions of the defect-free FSWed specimen, especially in the nugget zone, the grains are found to be completely transformed into fine-sized, uniformly distributed, equally spaced grain structures. This transformation of grains has taken place, because of the ideal up-roaring action and force of tool pin geometry, which has generated recrystallization of grains in a dynamic manner [34] and [35]. Based on the careful observation of these simulated temperature contour graphs, we can infer that performing the joining of AZ80A Mg alloys by Experimental Identification of Optimized Process Parameters for FSW of AZ91C Mg Alloy Using Quadratic Regression Models 741 Strojniski vestnik - Journal of Mechanical Engineering 66(2020)12, 736-751 (t>)> 0 By XO o o/e O 1« Prtdicttd Vidd (C) S.o I | 4 0 l e oo o 0/ y X O O / 0 0 Predlcttil tloji(9tkin Fig. 3. Graphical representation of the predicted values of the responses vs actual experimental values for a) actual vs. predicted ultimate tensile strength; b) actual vs. predicted yield strength and c) actual vs. predicted percentage of elongation FSW, at an optimized combination of higher rotational speeds of FSW tool, with the FSW tool traversing at low speeds and by applying larger values of axial load, will result in the generation of the ideal peak temperature, which will eventually contribute to perfect bonding between the AZ80A Mg alloy plates to be welded, thereby resulting in sound quality weldments. In addition to this, the employed taper cylindrical pin profile also has contributed to a greater extent in fabricating these sound quality weldments. This is because tapered tool pin profiles are always found to exhibit eccentricity. This eccentricity allows the smooth flow of the plasticized material around the tool pin profile region, thereby leading to the generation of a efficient stirring action, thereby contributing towards homogenous, finely spaced and refined grain structures in the zone of nugget [36] and [37]. 3.2 Impact on Mechanical Properties Fig. 6 illustrates the three-dimensional (3D) response surface charts for the tensile strength, plotted using the developed mathematical models. The maximum values of the responses are indicated by the apex of the surfaces of responses. From these graphs, the impingement of the several input parameters on the mechanical strength of the welded specimen can be easily visualized. For example, from Fig. 6c, it is visible that the mechanical strength of the welded joints has increased with the simultaneous increase in the tool traversing speed and rotational speed of the tool. It can also be visualized that the higher the rotating speed of FSW tool is, the larger is the tensile strength of the welded joint. From these graphs, it can be understood that the ultimate strength of tensile is lower at low rotating speeds (500 rpm, 650 rpm, etc.) and reaches its maximum value at higher speeds of rotational speed (1000 rpm to 1100 rpm). Similarly, greater values of tensile strength are exhibited at higher traversing speed of the FSW tool (1.0 mm/s to 1.5 mm/s). At the same time, the moderate amount of axial force (4 kN to 5 kN) and the tool with taper cylindrical pin geometry are found to display higher values of tensile strength in combination with the above-mentioned process parameters. Fig. 7 illustrates the 3D response surface charts for other mechanical properties obtained using the developed mathematical models. From these graphs, the impingement of several input parameters on the yield strength and percentage of elongation of the fabricated joints can be visualized easily. From these graphs, it is evident that the tool with taper cylindrical pin geometry exhibits defect-less AZ91C joints. This is mainly due to the ideal up-roaring action and the force of tool's taper cylindrical pin geometry, which 742 Satheesh, C. - Sevvel, P. - Senthil Kumar, R. Strojniški vestnik - Journal of Mechanical Engineering 66(2020)12, 736-751 (c) (b) (b) (c) (a) <"> Fig. 4. Optical macrograph of the cross-section of the fabricated high strength tensile specimen, indicating the various zones of the friction stir welded region namely; a) nugget region; b) thermo-mechanically transformed region; c) heat-affected region, and d) unaffected parent metal Fig. 5. Microstructures of a) AZ91C Mg alloy and various portions of the defect-free joint including, b) AZ91Cportion influenced by the up roaring force of tool shoulder, c) heat-affected portion, d) thermo-mechanically affected portion, and e) centre of the nugget zone (d) have caused the grains to undergo severe plastic deformation. The combination of this tapered pin geometry along with a high tool rotating speed (1000 rpm to 1100 rpm) have generated an ideal volume of heat, leading to an appreciable amount of plastic shear deformation, which has transformed the unequally sized grains of parent metal into uniaxial, radical, homogeneous uniformly distributed and equispaced fine grains, as seen in Fig. 6. Likewise, from the 3D response surface plots, we can understand that the process parameters (i.e., axial force and the tool's traversing speed) also play a vital role in deciding the strength of the welded joints. It is seen from these graphs that, the material flow in the zone of the nugget of the fabricated joints are dominated by the purge process during which the applied amount of force (axial) and the traversing speed of tool together push the plasticized substances, after the occurrence of severe plastic deformation. The employment of lower values of axial force (2 kN and 3 kN) along with lower tool traversing speeds (0.5 mm/s to 1.25 mm/s) have resulted in the generation of low peak temperatures in the nugget regions leading to incomplete dynamic recrystallization. This inappropriate combination has resulted in the fabrication of AZ91C joints with various defects leading to lower tensile strength values (189 kN to 207 kN) and lower yield strength values (100 kN to 114 kN). When an sufficient amount of axial force (4 kN to 5 kN) is applied during FSW of AZ91C joints on taper cylindrical pin profiled tool, along with suitable combination of 1000 rpm to 1100 rpm (rotating Experimental Identification of Optimized Process Parameters for FSW of AZ91C Mg Alloy Using Quadratic Regression Models 743 Strojniski vestnik - Journal of Mechanical Engineering 66(2020)12, 736-751 Fig. 6. 3D response charts illustrating the impact of a) axial force and tool traversing speed, b) axial force and rotating speed, c) tool traverse speed and rotating speed, d) tool pin profile and axial force, e) geometry of tool pin and rotational speed, and f) tool pin geometry and tool traverse speed on Tensile strength of welded joints speed), by making the FSW tool to travel in the range of 1.3 mm/s to 1.5 mm/s, the fictional heat was found to be generated in ideal and sufficient volumes, which directly contributes to the appreciable rise in the heat input, leading to preferable amount of plastic deformation, thereby accelerating the grain growth, 744 Satheesh, C. - Sevvel, P. - Senthil Kumar, R. Strojniški vestnik - Journal of Mechanical Engineering 66(2020)12, 736-751 Fig. 7. 3D response charts illustrating the impact of a) axial force and tool traversing speed, b) axial force and rotating speed, c) tool traverse speed and rotating speed on Yield strength, d) tool pin profile and axial force, e) geometry of tool pin and rotational speed, and f) tool pin geometry and tool traverse speed on percentage of elongation causing the transformation of unequally, large-sized grains into uniaxial, radical, homogeneous uniformly distributed and equispaced fine grains. 3.3 Analysis of SEM Images By observing Fig. 8a, we can visualize the presence of massive sized, coarse and irregularly distributed Experimental Identification of Optimized Process Parameters for FSW of AZ91C Mg Alloy Using Quadratic Regression Models 745 Strojniski vestnik - Journal of Mechanical Engineering 66(2020)12, 736-751 Fig. 8. SEM images of a) AZ91C Mg alloy, b) first interface region of the base metal and the nugget zone, c) thermo mechanically transformed region, and d) center of the nugget region grains containing Al12Mg17 intermetallic type aggregates. Concurrently, the observation of the scanning electron micrograph (SEM) pictures of the successfully fabricated AZ91C weldments (especially heat affected zone), as seen in Fig. 8b, reveals that the stirring process and force exerted by the taper cylindrical pin profiled tool have completely deliquesced the aggregates of Al12Mg17, and the ideal volume of generated heat (resulting from the suitable combination of the above mentioned FSW process parameters) have directed the plastically deformed grains along the axes of rotation. The Al12Mg17 precipitates (white particles) have been found to be elongated and fragmented. Fig. 8c shows the thermos mechanically transformed region, where the grains of the parent metal have changed their direction, the precipitates are fragmented, and it is completely influenced by up-roaring action and force of tool pin geometry. Likewise, as seen in Fig. 8d, the grain matrices of the AZ91C Mg alloy have also been found to undergone complete plastic deformation exactly along the FSW tool axis and the nugget zone is found to possess fine-sized, homogeneous, uniformly distributed grains [38]. Energy-dispersive X-ray spectroscopy (EDX) images of the centre of nugget zone of the fabricated weldment (as illustrated as SEM image in Fig. 8d) is illustrated in the Fig. 9 and from this figure, the presence of intermetallics can be understood clearly. 3.4 Determination of Optimized Process Parameters By analysing and visualizing these 3D response surface plots and the SEM images of the defect-free fabricated weldments, it is evident that the collaborative and collective impact of the axial force, traversing speed and rotating speed of tool have played a vital role in improving the tensile strength, resulting in AZ91C welded specimen, which are experimentally proven to be flawless. The various anticipated values of the mechanical properties are determined using 746 Satheesh, C. - Sevvel, P. - Senthil Kumar, R. Strojniški vestnik - Journal of Mechanical Engineering 66(2020)12, 736-751 Fig. 9. EDX images of the center of nugget zone of the fabriated weldment Fig. 10. Ideal Plots for the mechanical properties of welded AZ91C specimen Minitab analytical software, and Fig. 10 illustrates the established ideal plots. Table 5 describes in detail the predicted optimum input FSW process parameters required for obtaining good quality welded joints, including the type of pin profiled tool to be employed, the speed at which that tool must be rotated during welding, the amount of axial force to be exerted on the workpiece and the rate at which that FSW tool must travel over the workpiece. In other words, for obtaining an improved and preferable tensile strength of 238.34 MPa in the fabricated AZ91CMg alloy joints, it is predicted using the 3D graphs, that, a tool with taper cylindrical pin geometry must be employed at 1045.46 rpm, it should travel over the AZ91C specimen at 1.5 mm/s, employing a 4.87 kN force (axially). Under the above-mentioned combination of FSW process parameters, it is predicted that the fabricated joints would exhibit a yield strength of 142.58 kN, together with a percentage of elongation of 5.69. Six AZ91C Mg alloy welded specimen were fabricated using input parameter combinations described in Table 6. The obtained welded joints were free from flaws and exhibited mechanical properties, nearer to the predicted values, which indirectly reveals 747 Experimental Identification of Optimized Process Parameters for FSW of AZ91C Mg Alloy Using Quadratic Regression Models 65 Strojniski vestnik - Journal of Mechanical Engineering 66(2020)12, 736-751 Table 5. Description of the predicted optimum factors of FSW process Tool pin profile_Rotational speed [rpm] Tool traversing speed [mm/s] Axial force [kN] UTS [MPa] Yield [MPa] Elongation [%] Taper cylindrical 1045.46 1.5 4.87 238.34 142.58 5.69 Table 6. Comparison of the confirmatory experimental results with that of the results of the predicted values FSW Parameter_UTS_Yield Strength_% Elongation Trial run P NR Nt F Exp Predicted % of Error Exp Predicted % of Error Exp Predicted % of Error 1 1 -1 0 -1 218 209.55 -3.88 124 121.19 -2.27 3.94 4.25 7.87 2 2 1 2 1 232 231.27 -0.31 137 136.45 -0.40 5.63 5.89 4.62 3 1 2 -1 2 217 221.27 1.97 131 129.08 -1.47 5.61 5.88 4.81 4 0 1 2 0 225 223.25 -0.78 129 127.04 -1.52 4.62 4.42 -4.33 5 -1 2 -1 0 183 185.87 1.57 90 92.64 2.93 4.23 4.12 -2.60 6 Taper cylindrical 1045.46 rpm 1.5 mm/s 4.87 kN 239 238.34 -0.28 142 142.58 0.41 5.55 5.69 2.52 Fig. 11. Direct impacts of a) tool pin geometry, b) rotating speed, c) tool traversing speed, and d) axial force on mechanical strength of welded joints of AZ91C Mg alloy their precise conformity with the anticipated software values. 3.4 Analysis of Direct impact of Process Parameters Fig. 11 illustrates the direct impacts of those input FSW parameters on the mechanical strength of the 748 AZ91C Mg alloy welded specimen. The mechanical strength of the welded joints increases directly with the rise in tool traversing speed, axial force, and rotating speed. Then, they are observed to decline, once they reach their optimum high values. The graphs reveal that the axial force and rotating speed have a significant part in influencing the strength of Satheesh, C. - Sevvel, P. - Senthil Kumar, R. Strojniški vestnik - Journal of Mechanical Engineering 66(2020)12, 736-751 the welded joints of AZ91C Mg alloy, along with the tool pin geometry. The tool with taper cylindrical pin geometry was observed to exhibit the highest value of mechanical strength and, at the same time, it can be seen that the straight square pin profiled FSW tool exhibits minimum values of the ultimate tensile strength. This is mainly because geometry of the tool pin plays a vital role in deciding the shear stress, thereby contributing directly for the frictional heat generation within the region of the weld (joint region). When the consolidation of the plasticized material occurs layer by layer, due to lack of sufficient volume of heat, it drastically reduces the joint strength. Likewise, both lower and higher speeds of tool rotation as well as tool traversing speeds have resulted in generation of lower and higher volumes of frictional heat respectively, thereby leading to phenomenon of poor bonding. From these facts, it is obvious that perfect and suitable volume of heat (due to friction) gets generated only at optimized values of rotating speed of tool (1045 rpm), 1.5 mm/s traversing speed, by employing a tool with taper cylindrical pin geometry and by exerting a 4.87 kN axial force. This ideal volume of generated heat facilitates the flow of plasticized material in a uniform manner, enables the diffusion of the materials in the weld regions in a preferable manner, recrystallizes them perfectly, thereby producing defect-free AZ91C Mg alloy joints. 3 CONCLUSIONS This experimental investigation and analysis were carried out to determine the optimized values for various friction stir welding process parameters, essential for fabricating producing defect-free AZ91C Mg alloy joints. The various parameters of the FSW technique taken into during this investigational analysis include the rotating tool speed, the axial force applied, tool traversing speed and the tool's pin geometry. The obtained experimental results and observations from this investigation are listed below: • Regression quadratic equations and empirical relationships were established to determine and predict the responses (i.e., mechanical properties) for AZ91C Mg alloy joints fabricated using FSW technique. • The established quadratic regression model was found to be efficient in determining the responses of the FSWed AZ91C Mg allot joints in the range of +10 % of their actual experimental values at a 95 % level of confidence, and experimental trial runs were performed to confirm the predicted results. • 3D response surface charts illustrating the influence of input parameters on the mechanical properties and the direct impact plots representing the individual impact of these parameters on mechanical strength were generated and analysed in detail. • The tool with taper cylindrical pin geometry fabricated joints with larger tensile strength values leading to high-quality defect-free AZ91C Mg alloy joints. This was mainly due to the generation of an ideal volume of heat generation, unique pulsating effect and thereby leading to enhanced flow of the plasticized material uniformly. • AZ91C Mg alloy joints obtained using the tool having taper cylindrical pin geometry, employed at 1045 rpm, along a 1.5 mm/s tool travelling speed, under the exertion of a 4.87 kN axial force, were found to be free from flaws and naturally exhibited exceptional mechanical properties, especially higher values of ultimate tensile strength. • The investigational study of the micro and SEM images revealed that the grains in the zone of nugget have been completely remodelled into homogenous, uniformly distributed, refined grains, which in turn contributed to the exceptional mechanical properties. • Six AZ91C Mg alloy joints were fabricated employing the generated optimized input parameters. 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Experimental Identification of Optimized Process Parameters for FSW of AZ91C Mg Alloy Using Quadratic Regression Models 751 Strojniški vestnik - Journal of Mechanical Engineering 66(2020)12 Vsebina Vsebina Strojniški vestnik - Journal of Mechanical Engineering letnik 66, (2020), številka 12 Ljubljana, december 2020 ISSN 0039-2480 Izhaja mesečno Razširjeni povzetki (extended abstracts) Gorazd Lojen, Janez Mayer, Tonica Bončina, Franc Zupanič: Enostopenjska toplotna obdelava za obnovo mehanskih lastnosti hladno deformiranega jekla za rudniške podpore 31Mn4 SI 87 Wending Li, Guanglin Shi, Chun Zhao, Hongyu Liu, Junyong Fu: Vodenje elektrohidravličnega aktuatorja z nevronsko mrežo RBF po metodi drsnega režima s sestopom SI 88 Jianyong Liu, Jianfei Sun, Uzair Khaleeq uz Zaman, Wuyi Chen: Vpliv obrabe in geometrije orodja na drdranje, rezalne sile in integriteto površine titanove zlitine TB6 pri obdelavi s trdokovinskimi rezkarji različnih geometrij SI 89 Xibing Li, Zhe Yu, Xizhao Li, Weixiang Li, Tengyue Zou: Plužno-ekstruzijsko preoblikovanje za izdelavo toplotnih cevi z mikroutori na hidrostatičnem aksialnem drsnem ležaju težkega obdelovalnega stroja SI 90 C Satheesh, P Sevvel, R Senthil Kumar: Eksperimentalna identifikacija optimalnih parametrov procesa varjenja magnezijeve zlitine AZ91C z gnetenjem s kvadratnimi regresijskimi modeli SI 91 Strojniški vestnik - Journal of Mechanical Engineering 66(2020)12, SI 87 © 2020 Journal of Mechanical Engineering. All rights reserved. Prejeto v recenzijo: 2020-06-16 Prejeto popravljeno: 2020-09-09 Odobreno za objavo: 2020-11-19 Enostopenjska toplotna obdelava za obnovo mehanskih lastnosti hladno deformiranega jekla za rudniške podpore 31Mn4 Gorazd Lojen1* - Janez Mayer2 - Tonica Bončina1 - Franc Zupanič1 JUniverza v Mariboru, Fakulteta za strojništvo, Slovenija 2Premogovnik Velenje, d. d., Slovenija Jeklo 31Mn4 se pogosto uporablja za ločne rudniške podpore, tako v normaliziranem, vse pogosteje pa tudi v poboljšanem stanju + QT630 po DIN 21530-3. Za poboljšano stanje standard zahteva pri sobni temperaturi napetost tečenja Rp02 > 630 MPa, natezno trdnost Rm > 790 MPa, razteznost A5 > 16 % in udarno delo z DVM preizkušanci > 75 J (75 J z DVM preizkušanci je približno enakovredno 80 J z ISO-V preizkušanci). Najpomembnejša prednost ločnih podpor, sestavljenih iz segmentov iz profilov v obliki črke U je, da lahko segmenti loka pri prevelikem pritisku stene rova zdrsnejo eden po drugem. Pri tem se dodatno upognejo, prerez podpornega loka in pritisk nanj pa se zmanjšata. Tako se prepreči preobremenitev loka in podpora se ne poruši temveč ostaja funkcionalna, dokler je prerez rova dovolj velik ali dokler sposobnost drsenja segmentov ni izčrpana. Po demontaži se segmenti razdelijo v dve skupini: preveč poškodovani se takoj zavržejo, manj in dovolj enakomerno deformirani pa se lahko poravnajo in prevaljajo na ustrezen radij loka za ponovno vgradnjo. Hladne deformacije pri zdrsih segmentov med obratovanjem, ravnanjem in ponovnim valjanjem na potreben radij upogiba segmenta, utrjajo material in prej ali kasneje povzročijo prekomerno zmanjšanje duktilnosti. Zato se trenutno zavrže tudi precejšen delež tistih ločnih segmentov, ki bi jih bilo sicer še mogoče poravnati, vendar pa po popravilu ne bi več imeli zadostne duktilnosti. Trenutno v Premogovniku Velenje vsako leto zavržejo približno 3000 t segmentov jeklenih ločnih podpor. Ocenjujejo, da bi bilo približno 1/3, tj. okrog 1000 t, mogoče ponovno uporabiti, če bi bila na voljo zanesljiva in poceni toplotna obdelava, s katero bi jim povrnili ustrezno kombinacijo mehanskih lastnosti. V literaturi je mogoče najti precej poročil o raziskavah vplivov interkritičnega in subkritičnega žarjenja na sferoidizacijo perlita in povečanje duktilnosti različnih jekel. Poročil o raziskavah, povezanih z jeklom 31Mn4 je zelo malo, poročil o raziskavah, katerih cilj bi bil povrnitev natančno določene kombinacije mehanskih lastnosti hladno deformiranega poboljšanega jekla s podkritičnim žarjenjem, pa nismo našli, ne za jeklo 31Mn4 na za druga jekla. Zato je bil cilj te raziskave oceniti izvedljivost stroškovno učinkovite enostopenjske toplotne obdelave za obnovo zahtevane kombinacije mehanskih lastnosti prekomerno hladno utrjenih poboljšanih jeklenih segmentov ločnih podpor. Vzorci jekla 31Mn4 + QT630 so bili izrezani iz novega profila TH 29. Da smo ugotovili razmerje med stopnjo hladne deformacije in trdoto, so bili različno močno hladno deformirani v območju od 0 % do 45 %. Nato smo vzorce žarili od 30 min do 8 h pri temperaturah od 450 °C do 700 ° C, da bi ugotovili primerne kombinacije časa in temperature žarjenja. S svetlobnim in elektronskim vrstičnim mikroskopom so bile pregledane mikrostrukture nedeformiranih, hladno deformiranih in rekristaliziranih vzorcev. Za vsa stanja so bile narejene tudi meritve trdote, natezni preizkusi in preizkusi udarne žilavosti po Charpyju. Mikrostruktura nedeformiranega materiala ni bila značilna za poboljšana jekla. Bila je zelo nehomogena in je vsebovala znatne deleže predevtektoidnega ferita in perlita, ki so posledica neustrezno izvedene toplotne obdelave. Razteznost je bila nekoliko manjša od zahtev standarda, preostale mehanske lastnosti pa so bile ustrezne. Trdota je bila pod 250 HV 30. S hladno deformacijo se je trdota povečevala skoraj linearno, za približno 2 HV na 1 % hladne deformacije. To pomeni, da je mogoče enostavno in zanesljivo ugotoviti dejansko stopnjo hladne deformacije, če je poznana prvotna trdota. Po enourni rekristalizaciji pri temperaturah od 600 °C do 620 °C je bila mikrostruktura veliko bolj homogena. Sestavljena je bila iz feritne matrice, ki je vsebovala enakomerno razpršene globularne karbidne delce, kar je značilno za pravilno kaljena in popuščana jekla. Udarno delo KV pri Charpyjevem preizkusu ISO-V preizkušancev je bilo nad 90 J, podobno kot pred hladno deformacijo. Razteznost A5, napetost tečenja Rp02 in trdota po Vickersu HV pa so bile pri večini vzorcev večje kot pred hladno deformacijo. S tem so rezultati potrdili, da je možno z ustrezno toplotno obdelavo v enem koraku obnoviti oz. celo izboljšati prvotne mehanske lastnosti prekomerno hladno deformiranega jekla 31Mn3+QT630. Ker se trdnost, duktilnost in udarna žilavost med rekristalizacijo pri močneje deformiranem materialu spreminjata hitreje kot pri manj deformiranem, je mogoče parametre toplotne obdelave optimizirati tako, da se z istimi parametri povrnejo ustrezne lastnosti različno močno deformiranemu materialu. Z izvajanjem protokola obdelave, ki zajema določitev stopnje hladne deformacije in ustrezno enostopenjsko toplotno obdelavo, bi bilo mogoče povečati stopnjo ponovne uporabe segmentov jeklenih lokov, obenem pa zmanjšati stroške za nove profile in posredno zmanjšati vplive rudarjenja na okolje. Ključne besede: mikrostruktura, deformacijsko utrjanje, mehanske lastnosti, rekristalizacija, jeklo 31Mn4, jeklo 1.0520, jekleni nosilci *Naslov avtorja za dopisovanje: Univerza v Mariboru, Fakulteta za strojništvo, Smetanova ulica 17, 2000 Maribor, Slovenija, gorazd.lojen@um.si SI 87 Strojniški vestnik - Journal of Mechanical Engineering 66(2020)12, SI 88 © 2020 Strojniški vestnik. Vse pravice pridržane. Prejeto v recenzijo: 2020-07-19 Prejeto popravljeno: 2020-10-19 Odobreno za objavo: 2020-11-06 Vodenje elektrohidravličnega aktuatorja z nevronsko mrežo RBF po metodi drsnega režima s sestopom Wending Li123* - Guanglin Shi1 - Chun Zhao23 - Hongyu Liu23 - Junyong Fu23 JSola za strojništvo, Univerza v Šanghaju Jiao Tong, Kitajska 2Šanghajski inštitut za regulacijsko tehniko v letalstvu, Kitajska 3Šanghajsko tehnično raziskovalno središče za servosisteme, Kitajska Cilj predstavljenega dela je razrešitev problema nelinearnega vodenja elektrohidravličnega aktuatorja in v ta namen je bil postavljen nelinearen model sistema. V raziskavi nove strategije vodenja sta bili opredeljeni stabilnost in natančnost vodenja pri velikih motilnih obremenitvah, regulacijski zakon pa je bil razrešen za neznane in negotove pogoje. Z uvedbo opazovalca motenj in v kombinaciji z nevronsko mrežo za vodenje po metodi drsnega režima so bili odpravljeni vplivi neznanega modela in nemodelirane dinamike, razrešen je bil problem nelinearnega vodenja elektrohidravličnega sistema in izboljšana je bila natančnost vodenja. Za izboljšanje obstojnosti proti velikim motnjam in visoko natančnost nelinearnega aktuatorja so najprej določene nelinearne enačbe stanja elektrohidravličnega aktuatorja, ki vključujejo motnje. Tri enačbe stanja vsebujejo dva perturbacijska člena za uporabo sestopne metode pri reševanju problema perturbacij nelinearnega sistema. Velika prva motnja predstavlja obremenitev in je vključena z opazovalnikom motenj ter odpravljena z zaviralnim členom drsnega režima. Druga motnja je nemodelirana dinamika in je odpravljena z zaviralnimi členi drsnega režima. Neznani člen v nelinearnem modelu sistema je aproksimiran z adaptivnim algoritmom nevronske mreže. Pri konstruiranju enačbe napak sistema je bila uporabljena posebna metoda, ki razreši odvedljivost napake in odpravi »eksplozijo« odvoda. Virtualni regulacijski zakon in sistemski regulacijski zakon sta izračunana po sestopni metodi. Funkcija predznaka v virtualnem regulacijskem zakonu je zamenjana s posebno funkcijo na osnovi arkus tangensa za odpravo težav zaradi singularnosti. V raziskavi so uporabljene naslednje glavne metode: 1. Postavljen je nelinearen model elektrohidravličnega aktuatorja tretjega reda z motnjo; 2. Za odpravo različnih motenj sta uporabljena opazovalnik stanj in člen drsnega režima; 3. Z uvedbo sestopne metode je najprej razrešen virtualni regulacijski zakon, nato pa je razrešen še sistemski regulacijski zakon z nevronsko mrežo za vodenje po metodi drsnega režima, ki dosega visoko stopnjo natančnosti; 4. Opravljene so bile simulacije delovanja pri visokih in nizkih hitrostih, primerjava rezultatov eksperimentov z rezultati simulacij pa je pokazala dobro ujemanje. V enačbo stanja nelinearnega sistema sta bila vključena dva perturbacijska člena, ki zagotavljata dobre pogoje za omejevanje perturbacij. Prva motnja predstavlja velike obremenitve, druga pa nemodelirano dinamiko sistema. Uvedba dveh različnih členov za motnje zagotavlja večjo prilagodljivost pri načrtovanju strategije vodenja in izbiro ustreznega pristopa za omejevanje različnih motenj. Sestopna metoda ima lahko zelo pomembno vlogo pri načrtovanju strategije vodenja nelinearnih sistemov. Najprej je bil določen virtualni regulacijski zakon, nato pa je bil izračunan regulacijski zakon sistema. Uporabiti je mogoče različne metode načrtovanja za optimalno zmogljivost sistema. Uporaba drsnega režima za omejevanje napak in motenj zahteva uvedbo simbolnih funkcij. Virtualni regulacijski zakon vsebuje funkcijo predznaka, ki je ni mogoče odvajati. Za aproksimacijo te funkcije je bila zato uporabljena posebna funkcija z arkus tangensom. Pri modeliranju nelinearnega sistema ni mogoče določiti natančnega modela in potrebna je aproksimacija z nevronsko mrežo RBF. Napaka aproksimacije nevronske mreže RBF je odpravljena z zaviralnim členom drsnega režima. S kombiniranim pametnim algoritmom je bolje razrešen problem strategije vodenja nelinearnega sistema. V simulacijah in eksperimentih sta bila zajeta dva značilna seta delovnih pogojev, ki dobro predstavljata delovne pogoje predmeta raziskav. Rezultati simulacij in analiz dokazujejo reprezentančnost delovnih pogojev. V predstavljeni študiji nelinearne strategije vodenja je omejevanje zunanje motnje izvedeno z opazovalnikom, brez uvedbe razlike tlakov v komorah hidravličnega valja. V prihodnjih raziskavah bo opazovalnik motenj mogoče nadgraditi z uvedbo omenjene tlačne razlike za izboljšanje natančnosti vodenja. Uporabljen je bil inovativen pristop k načrtovanju strategije vodenja nelinearnega sistema, ki povezuje sestopno metodo in nevronsko mrežo po metodi drsnega režima. Novost v sestopni metodi sta uporaba aproksimacije za funkcijo predznaka in uvedba opazovalnika motenj ter. uporaba različnih strategij za razrešitev motnje. V kombinaciji z nevronsko mrežo RBF in strategijo drsnega režima so dobro razrešene nelinearnost, motnje in negotovost modela sistema. Natančnost vodenja in odpravljanje motenj pri delovanju elektrohidravličnega aktuatorja sta bila izboljšana s sestopno metodo. Simulacije in eksperimenti kažejo dobro in robustno delovanje strategije vodenja. Ključne besede: nevronska mreža RBF, drsni režim, sestop, nelinearno vodenje, elektrohidravlični aktuator, opazovanje motenj, regulacijski zakon SI 88 *Naslov avtorja za dopisovanje: Šola za strojništvo, Univerza v Šanghaju Jiao Tong, 800 Dongchuan Road, Minhang District, Shanghai 200240, Kitajska, liwending1983@163.com Strojniški vestnik - Journal of Mechanical Engineering 66(2020)12, SI 89 © 2020 Strojniški vestnik. Vse pravice pridržane. Prejeto v recenzijo: 2020-04-10 Prejeto popravljeno: 2020-06-22 Odobreno za objavo: 2020-07-06 Vpliv obrabe in geometrije orodja na drdranje, rezalne sile in integriteto površine titanove zlitine TB6 pri obdelavi s trdokovinskimi rezkarji različnih geometrij Jianyong Liu1 - Jianfei Sun2 3 4 - Uzair Khaleeq uz Zaman5* - Wuyi Chen23 1 AECC Shenyang Liming Aero-Engine Co., Ltd., Kitajska 2 Univerza Beihang, Šola za strojništvo in avtomatizacijo, Kitajska 3 Sodelovalno inovacijsko središče za napredne letalske motorje, Kitajska 4 Tehnološko raziskovalno središče v Pekingu za visokoučinkovite in zelene CNC-obdelovalne procese in opremo, Kitajska 5 Nacionalna znanstveno-tehniška univerza, Kolidž za elektrotehniko in strojništvo, Pakistan V predstavljenem delu so bili uporabljeni rezkarji za obdelavo brez vibracij (steblasti rezkarji z variabilno vijačnico) (VH) in variabilnim korakom (VP)) in standardni (SP) steblasti rezkarji za obdelavo titanove zlitine TB6 (Ti-10V-2Fe-3Al) s ciljem preučitve vpliva parametrov obrabe in geometrije orodja na drdranje, rezalne sile in integriteto obdelanih površin. Titanova zlitina TB6 je zelo razširjena v letalski industriji, kjer se uporablja za izdelavo letalskih trupov in kril, pristajalnih mehanizmov in delov helikopterskih rotorjev. Razlog za to je v njenih lastnostih, kot so visoka specifična trdnost, odlična obstojnost proti koroziji in dobra odpornost proti utrujanju. Nekatere druge lastnosti zlitine TB6, kot so nizka toplotna prevodnost, majhen modul elastičnosti in visoka kemična aktivnost, pa lahko povzročijo tudi visoke temperature pri odrezavanju, velike rezalne sile, krajšo obstojnost orodja, nizko stopnjo odvzema materiala in neustrezno kakovost površine obdelovancev. Izboljšanje integritete površine obdelovancev iz titanove zlitine TB6 predstavlja poseben izziv zaradi njene neugodne obdelovalnosti. Integriteta površin je pomembna tudi za samo delovanje, saj sta od nje odvisni funkcionalnost in trajna nihajna trdnost kritičnih komponent. Integriteta površin pomembno vpliva na kakovost površin in je odvisna od različnih dejavnikov, kot so vibracije, rezalni parametri, obraba orodja, geometrijski parametri orodja itd. Mnogi raziskovalci poročajo o tem, da geometrijski parametri rezkarjev pomembno vplivajo na kakovost obdelanih površin. Zlasti pri orodjih z variabilno vijačnico in variabilnim korakom so se izvajale raziskave za izboljšanje stabilnosti in zmanjšanje drdranja s ciljem izboljšanja kakovosti rezkanja. Objavljena literatura pa se le malo ukvarja z ostalimi kazalniki kakovosti površin po obdelavi z rezkarji VH in VP, kot je tekstura površine, ki nastane zaradi drdranja. Glavni prispevek članka je najnovejši opis vpliva obrabe, drdranja in rezalnih sil na integriteto površine delov iz titanove zlitine TB6 po obdelavi s trdokovinskimi rezkarji različnih geometrij. Novost je v uporabljenih eksperimentalnih tehnikah oz. v analizi vplivov obrabe, drdranja in rezalnih sil na integriteto površine titanove zlitine TB6 po obdelavi. Glavni rezultati analize so: i. Geometrija orodja pomembno vpliva na obrabo rezkarjev. V primerjavi z rezkarji SD se rezkarji VP in VH s spremenjeno vijačnico in korakom obrabljajo hitreje, sama obraba pa ni enakomerna. Ta neenakomernost se zmanjšuje z naraščanjem obsega obrabe. ii. Obraba rezkarjev VP, VH in SD pomembno vpliva na drdranje in na rezalne sile. Drdranje in rezalne sile so pri vseh treh vrstah rezkarjev naraščali s povečanjem obrabe. iii. Sprememba geometrije orodja je učinkovito sredstvo za omejevanje drdranja. S spremembo koraka in kota vijačnice pri rezkarjih VP in VH je mogoče zmanjšati drdranje v primerjavi s standardnimi rezkarji. iv. Geometrija in obraba orodja vplivata na sile med obdelavo z rezkarji VP in VH. Čeprav sprememba vijačnice in kota vpliva na velikost rezalnih sil v primerjavi s standardnimi rezkarji, obraba zmanjša radialno globino rezanja in poveča ekstruzijski pritisk ter tako vpliva na rezalne sile med obdelavo s temi rezkarji. v. Površinska hrapavost se je povečala s povečanjem obrabe treh vrst rezkarjev, hrapavost po obdelavi z rezkarji VH in VP pa je bila manjša kot po obdelavi z rezkarji SD. Z naraščanjem obrabe so pri vseh rezkarjih glavni vir napetosti preostale tlačne napetosti. Mikrotrdota se je pri vseh rezkarjih povečala s povečanjem obrabe v trdem predelu in v predelu jedra, zmanjšala pa se je s povečanjem obrabe v mehkejšem predelu. Ključne besede: drdranje, rezalne sile, integriteta površin, titanova zlitina TB6, obraba *Naslov avtorja za dopisovanje: Nacionalna znanstveno-tehniška univerza, Kolidž za elektrotehniko in strojništvo, Islamabad, 44000, Pakistan, uzair.khaleeq@ceme.nust.edu.pk SI 89 Strojniški vestnik - Journal of Mechanical Engineering 66(2020)12, SI 90 © 2020 Strojniški vestnik. Vse pravice pridržane. Prejeto v recenzijo: 2020-08-25 Prejeto popravljeno: 2020-11-20 Odobreno za objavo: 2020-11-24 Plužno-ekstruzijsko preoblikovanje za izdelavo toplotnih cevi z mikroutori na hidrostatičnem aksialnem drsnem ležaju težkega obdelovalnega stroja Xibing Li1 - Zhe Yu - Xizhao Li2 - Weixiang Li1 - Tengyue Zou1* 1 Univerza za kmetijstvo in gozdarstvo province Fujian, Kolidž za strojništvo in elektrotehniko, Kitajska 2 Poklicni in tehnični kolidž province Shandong, Oddelek za pametno proizvodnjo, Kitajska Hidrostatični aksialni drsni ležaj je ključna komponenta težke vertikalne CNC stružnice. Zaradi pomanjkljivega odvajanja toplote med procesom obdelave lahko nastopijo težave z mazanjem in posledično upad kakovosti obdelave. Možna rešitev je uporaba toplotnih cevi z mikroutori za izboljšano odvajanje toplote iz komponent. Za izdelavo toplotnih cevi z mikroutori na hidrostatičnem aksialnem drsnem ležaju je predlagan postopek plužno-ekstruzijskega preoblikovanja, ki zagotavlja boljše lastnosti kot tradicionalna elektroerozijska obdelava. Plužno-ekstruzijsko preoblikovanje kovin ima lastnosti oranja in ekstrudiranja in se izvaja s posebnim večzobim orodjem na obdelovalnem stroju, ki plastično deformira del kovinske stene in oblikuje rebra z veliko mikroutori. Izdelana so bila različna večzoba orodja za plužno-ekstruzijsko preoblikovanje iz ležajnega jekla Gcr15 in hitroreznega jekla W18Cr4V V predstavljenem delu je ovrednotenih šest večzobih orodij z različnimi konstrukcijskimi parametri. Ugotovljeno je bilo, da je med njimi najprimernejše orodje z dolžino 10 mm, 55 zobmi, globino zob 0,25 mm in zgornjo širino zob 0,22 mm. Največja kakovost površine mikroutorov je bila ugotovljena pri eksperimentih s hitrostjo preoblikovanja 50 mm/min. Zaradi prevelike hitrosti preoblikovanja se lahko poškoduje vlečni drog, če je hitrost preoblikovanja premajhna, pa ni zaželenega učinka oranja. Toplotna cev ima sicer odlično toplotno prevodnost, toda toplotne obremenitve ni mogoče povečevati v neskončnost zaradi različnih dejavnikov, ki omejujejo sposobnost prenosa toplote (meja viskoznosti, meja nosilnosti, meja vrenja, kapilarna meja itd.). Toplotna prevodnost cevi je odvisna predvsem od stanja na notranji steni, zato se bo nadaljevalo delo na optimizaciji parametrov cevi in obdelovalnih postopkov za postopno izboljševanje kapilarnega delovanja cevi. V prihodnjih raziskavah se bo mogoče posvetiti tudi izboljšanju preoblikovalnega postopka za bolj dovršeno izvedbo mikroutorov, povečanje različnih mejnih vrednosti in učinkovitejši prenos toplote. V preizkusih prenosa toplote je bil potrjen tudi teoretični model toplotne cevi z mikroutori, izdelane s plužno-ekstruzijskim preoblikovanjem. Članek predstavlja plužno-ekstruzijsko preoblikovanje za izdelavo toplotnih cevi z mikroutori, ki lahko izboljšajo disipacijo toplote pri težkih obdelovalnih strojih. Analiza digitalnih posnetkov je pokazala, da lahko ta postopek zagotavlja večjo zmogljivost in nižje stroške kot tradicionalna elektroerozijska obdelava. Podani so tudi priporočeni parametri večzobega orodja in hitrosti plužno-ekstruzijskega preoblikovanja. Članek bo uporaben za vse inženirje, ki se ukvarjajo s hlajenjem ali s fino obdelavo. Ključne besede: toplotna cev, disipacija toplote, mikroutor, težki stroji, hidrostatični aksialni drsni ležaj, plužno-ekstruzijsko preoblikovanje SI 90 *Naslov avtorja za dopisovanje: Univerza za kmetijstvo in gozdarstvo province Fujian, Kolidž za strojništvo in elektrotehniko, Fuzhou, Kitajska, zouty@fafu.edu.cn Strojniški vestnik - Journal of Mechanical Engineering 66(2020)12, SI 91 Prejeto v recenzijo: 2020-09-10 © 2020 Strojniški vestnik. Vse pravice pridržane. Received revised form: 2020-11-26 Accepted for publication: 2020-12-02 Eksperimentalna identifikacija optimalnih parametrov procesa varjenja magnezijeve zlitine AZ91C z gnetenjem s kvadratnimi regresijskimi modeli C. Satheesh1 - P. Sevvel2* - R. Senthil Kumar2 1 Oddelek za strojništvo, Inštitut za inženiring in tehnologijo Madha, Indija 2 Oddelek za strojništvo, Tehniški kolidž S. A., Indija AZ91C je zelo razširjena in priljubljena magnezijeva zlitina, ki je uporabljena za 85 % do 91 % vseh magnezijevih ulitkov zaradi svoje edinstvene livnosti, nizkih stroškov, visoke stopnje obstojnosti proti koroziji itd. Obenem pa obstaja tudi potreba po izboljšanju njene nizke razteznosti (zaradi grobih zrn v mikrostrukturi), s čimer bi zlitina postala uporabna tudi za druge industrijske sektorje. Izvajajo se različni raziskovalni programi za razširitev uporabnosti magnezijeve zlitine AZ91C, ki so osredotočeni predvsem na izboljšanje duktilnosti zlitine z zmanjševanjem kristalnih zrn. Izdelava trdnih zvarnih spojev s kakršno koli tehniko zahteva uporabo idealnih oz. optimalnih vhodnih parametrov. Vse sestave, ki se varijo konvencionalno z različnimi kombinacijami procesnih parametrov, je treba po izdelavi pregledati in potrditi, da izpolnjujejo pričakovane standarde. S konvencionalnimi parametri lahko nastanejo zvarni stroji, ki so blizu uporabnikovim pričakovanjem, ne dosegajo pa optimalne kombinacije procesnih parametrov. Predmet predstavljene eksperimentalne raziskave je napovedovanje idealne oz. optimalne kombinacije vhodnih parametrov pri varjenju z gnetenjem za doseganje najboljših mehanskih lastnosti, vključno s plastičnostjo, natezno trdnostjo in razteznostjo. V ta namen je bila uporabljena napredna statistična in matematična tehnika - metoda odzivnih površin. Oblikovan je bil kvadratni regresijski model, ki vključuje različne vhodne kriterije varjenja z gnetenjem za napovedovanje in določitev odgovora pri varjenju delov iz magnezijeve zlitine AZ91C. Upoštevani vhodni parametri so hitrost podajanja orodja, vrtilna hitrost orodja, profil (geometrija) čepa in aksialna sila. Uporabljen je bil štirifaktorski in petnivojski središčni sestavljeni načrt. Za razvoj 3D odzivnih površin za različne mehanske lastnosti po metodi RSM so bili oblikovani kvadratni regresijski modeli. Njihova primernost je bila preizkušena in validirana z analizo variance. Rezultati eksperimentov v obliki grafov raztrosa kažejo dobro ujemanje z modeli. Ugotovljene so bile izboljšane lastnosti zvarov delov iz magnezijeve zlitine AZ91C, ustvarjenih s prisekanim cilindričnim čepom ter vrtilno hitrostjo 1045 vrt/min, podajalno hitrostjo 1,5 mm/s in aksialno silo 4,87 kN. Izdelanih je bilo šest zvarov iz magnezijeve zlitine AZ91C z optimiziranimi vhodnimi parametri. Pregledanih je bilo vseh šest in ugotovljeno je bilo, da so zvari popolnoma brez napak. Ključne besede: magnezijeva zlitina AZ91C, kvadratni regresijski model, varjenje z gnetenjem, geometrija čepa, vrtilna hitrost orodja, porušitvena natezna trdnost, podajalna hitrost orodja, aksialna sila *Naslov avtorja za dopisovanje: Oddelek za strojništvo, Tehniški kolidž S. A., Tharapakkam, Chennai - 600 122, Indija, sevvel_ready@yahoo.co.in SI 91 Guide for Authors All manuscripts must be in English. Pages should be numbered sequentially. The manuscript should be composed in accordance with the Article Template given above. The maximum length of contributions is 12 pages (approx. 5000 words). Longer contributions will only be accepted if authors provide justification in a cover letter. For full instructions see the Information for Authors section on the journal's website: http://en.sv-jme.eu . SUBMISSION: Submission to SV-JME is made with the implicit understanding that neither the manuscript nor the essence of its content has been published previously either in whole or in part and that it is not being considered for publication elsewhere. All the listed authors should have agreed on the content and the corresponding (submitting) author is responsible for having ensured that this agreement has been reached. 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Strojniški vestnik -Journal of Mechanical Engineering Aškerčeva 6, 1000 Ljubljana, Slovenia, e-mail: info@sv-jme.eu http://www.sv-jme.eu Contents Papers Gorazd Lojen, Janez Mayer, Tonica Boncina, Franc Zupanic: Single-Step Heat Treatment for the Restoration of the Mechanical Properties of Cold-Strained Mining Support Steel 31Mn4 Wending Li, Guanglin Shi, Chun Zhao, Hongyu Liu, Junyong Fu: RBF Neural Network Sliding Mode Control Method Based on Backstepping for an Electro-hydraulic Actuator Jianyong Liu, Jianfei Sun, Uzair Khaleeq uz Zaman, Wuyi Chen: Influence of Wear and Tool Geometry on the Chatter, Cutting Force, and Surface Integrity of TB6 Titanium Alloy with Solid Carbide Cutters of Different Geometry Xibing Li, Zhe Yu, Xizhao Li, Weixiang Li, Tengyue Zou: Plough-extrusion Forming for Making Micro-groove Heat Pipes on Hydrostatic Thrust Bearings of Heavy Machinery Satheesh C, Sevvel P, Senthil Kumar R: Experimental Identification of Optimized Process Parameters for FSW of AZ91C Mg Alloy Using Quadratic Regression Models 687 697 709 9770039248001